Saimm 202110 oct

VOLUME 121

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OCTOBER 2021

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Nolitha Fakude President, Minerals Council South Africa Honorary Vice Presidents Gwede Mantashe Minister of Mineral Resources, South Africa Ebrahim Patel Minister of Trade, Industry and Competition, South Africa Blade Nzimande Minister of Higher Education, Science and Technology, South Africa President I.J. Geldenhuys President Elect Z. Botha Senior Vice President W.C. Joughin Junior Vice President E Matinde Incoming Junior Vice President G.R. Lane Immediate Past President V.G. Duke Honorary Treasurer W.C. Joughin Ordinary Members on Council Z. Fakhraei B. Genc K.M. Letsoalo S.B. Madolo F.T. Manyanga T.M. Mmola G. Njowa

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Editorial Board S.O. Bada R.D. Beck P. den Hoed I.M. Dikgwatlhe R. Dimitrakopolous* M. Dworzanowski* L. Falcon B. Genc R.T. Jones W.C. Joughin A.J. Kinghorn D.E.P. Klenam H.M. Lodewijks D.F. Malan R. Mitra* C. Musingwini S. Ndlovu P.N. Neingo M. Nicol* S.S. Nyoni N. Rampersad Q.G. Reynolds I. Robinson S.M. Rupprecht K.C. Sole A.J.S. Spearing* T.R. Stacey E. Topal* D. Tudor* F.D.L. Uahengo D. Vogt* *International Advisory Board members

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VOLUME 121 NO. 10 OCTOBER 2021

Contents Journal Comment

by S. Ndlovu. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . iv-v President’s Corner: When one tugs on a single thing in nature, one finds it attached to the rest of the world

by I.J. Geldenhuys . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . vi

ACID MINE DRAINAGE Sulphate removal technologies for the treatment of mineimpacted water M. van Rooyen, P.J. van Staden, and K.A. du Preez . . . . . . . . . . . . . . . . 523 This paper presents a high-level comparison of four sulphate removal technologies for remediating mine-impacted water including acid mine drainage (AMD). Each process is shown to be subject to a unique set of constraints which might favour one over another for a specific combination of location and water composition. The total cost calculated for each process also depends on the types of effluents that are allowed to be discharged.

Are pit lakes an environmentally sustainable closure option for opencast coal mines?

A.C. Johnstone. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 531 The aim of this study was to determine if pit lakes are a sustainable coal mine closure option in South Africa. The study investigated the water balance, chemistry, limnology, and bacteria of three selected pit lakes. The results indicated that pit lakes can be designed as ‘terminal sinks’ to provide a sustainable mine closure option.

A suggested design manual was developed to aid mine owners and regulators.

Directory of Open Access Journals

THE INSTITUTE, AS A BODY, IS NOT RESPONSIBLE FOR THE STATEMENTS AND OPINIONS ADVANCED IN ANY OF ITS PUBLICATIONS. Copyright© 2021 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 International Advisory Board 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 ofMcGill review University, or criticism under The Copyright Act no. 98, 1978, Section 12, of the R. Dimitrakopoulos, Canada 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 D. Dreisinger, University of British Columbia, Canada 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 M. Dworzanowski, Consulting Metallurgical Engineer, France contents of the publication without permission is always illegal.

U.S. Copyright Law applicable to users in the U.S.A. E. Esterhuizen, NIOSH Research Organization, USA H. Mitri, University, Canada The appearance of the statement of copyright at the bottom of theMcGill 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 ThisMurdoch consent isUniversity, given on condition that the copier pays the stated fee for each copy of a paper M.J. use.

Nicol, Australia 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, E. Topal, Curtin University, Australia 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 D. Vogt,works, University of Exeter, United Kingdom advertising or promotional purposes, for creating new collective or for resale.

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The effect of decarburization on the fatigue life of overhead line hardware J. Calitz, S. Kok, and D. Delport . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 537 Altering the microstructure in order to improve the tensile properties of bow shackles, resulted in inconsistency in the fatigue performance. This raises the question whether fatigue life can be attributed to microstructural changes along the profile of the shackle or to decarburization at the surface. Bow shackles forged from 080M40 (EN8) material were subjected to different heat treatments and to five different fatigue load cases.

Although the change in microstructure does improve both the tensile and fatigue performance, the depth of the decarburization layer has a greater effect on the high cycle fatigue life of bow shackles than the non-homogeneous microstructure.

Microstructure, microhardness, and tensile properties of hot-rolled Al6061/TiB2/ CeO2 hybrid composites S. Iyengar, D. Sethuram, R. Shobha, and P.G. Koppad . . . . . . . . . . . . . . . . . . . . . . . . . . 543 TiB2 and CeO2 particle-reinforced Al6061 hybrid composites were manufactured using stir casting and hot rolling techniques. The base alloy and composites were hot rolled at 500°C, and a 50% reduction was achieved through 12 passes. The effect of varying TiB2 and CeO2 particle additions on the microstructure and mechanical properties of the Al6061 matrix was studied. The optimum reinforcement combination was 2.5% CeO2 and 2.5% TiB2 particles.

Resistance spot welding of a thin 0.7 mm EN10130: DC04 material onto a thicker 2.4 mm 817M40 engineering steel K.A. Annan, R.C. Nkhoma, and S. Ngomane . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 549 The effects of welding current, electrode force, and welding time in resistance spot weld were studied to investigate the effectiveness of welded joints between a thin material and a thicker part through analysis of the microstructural and mechanical properties. All welded specimens were subjected to tensile testing at room temperature (25°C) and sub-zero temperature (–46°C), and the optimal welding conditons established. The welding current was found to be the most significant parameter influencing the quality of the weld joint.

Correlations of geotechnical monitoring data in open pit slope back-analysis – A mine case study A.F. Silva, J.M.G. Sotomayor, and V.F.N. Torres. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 556 Geotechnical monitoring plays an important role in the detection of operational safety issues in the slopes of open pits. This work consisted of an analysis of monitoring data (pore pressure and displacement) and its correlation with the calculated tension and displacement of the mass of an established failure slope using the finite element method. The results were consistent with both the measured displacements and the maximum deformation contours, revealing the possible failure mechanism, allowing the strength parameters to be calibrated according to the slope failure conditions, and providing information about the contribution of each variable to the failure event.

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nal

Jour

ment

Com

S

outh Africa has one of the most prominent mining industries in the world. The country saw a boom in the mining industry in the late 19th century with the drivers being gold and diamond mining, followed soon after by coal mining, then PGM processing. Today, gold, PGMs, and coal mining continue to make significant contributions to the economic and social development of the country. Despite the criticality of mining to the growth and development of South Africa and other nations across the globe, the industry is associated with process challenges and legacies of environmental impact, one of which, is the issue of mineimpacted water.

Mine-impacted water is considered to be one of the main pollutants of surface- and groundwater in many countries that have historical or current mining industries and its potential effects on natural resources, communities, and human health have become increasingly evident. Mine-impacted water has long been regarded as one of the most serious and pervasive challenges facing the mining and minerals industry.

While a wide range of technologies are being developed for preventing the generation of, and the control and remediation of, mine-impacted water, most of these approaches consider it a nuisance that needs to be quickly disposed of after minimum required treatment, in line with the legislation of that particular country. However, recently, there has been an emerging paradigm shift towards environmental responsibility and sustainable development. Thus, studies focusing on sustainable treatment technologies, value recovery from the waste solutions, mining closure practices, and legislation to mitigate potential future challenges arising from mine-impacted water have become predominant.

One of the best approaches to dealing with mine-impacted water is to consider it as a valuable resource and look at the recovery of clean water to satisfy the needs of a variety of mining and non-mining users. Since South Africa is a water-scarce country, this is a more practical and applicable approach to the problem. The production of other valuable and saleable by-products such as metals and salts that could be used to offset some of the operational costs is also being considered. In fact, recycling, and re-use of water and the recovery of value products is one of the emerging pragmatic approaches to mitigating the challenges associated with mine-impacted water.

It is at events such as conferences, workshops, and seminars that stakeholders can share unbiased, state-of-the-art expertise and knowledge, novel solutions and approaches, technical knowhow, and advocacy with respect to the legacy of, and sustainable solutions related to, mine-impacted water. Such events can help inspire and accelerate some of the work being done by all interested stakeholders on sustainable and holistic ways to deal with the issue of mine-impacted water. The papers in this edition of the Journal reflect some of the discussions arising from the conference held in November 2020.

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Journal Comment (continued) The conference, which was organized by the SAIMM in collaboration with the University of the Witwatersrand, Mintek in South Africa, and RWTH Aachen University in Germany, attracted speakers and authors from a number of countries such as South Africa, the UK, Germany, Nigeria, Zambia, Serbia, and Belgium. The idea of the conference was born from a collaborative project between Wits University through the School of Chemical and Metallurgical Engineering and the Institute IME Process Metallurgy and Metal Recycling at RWTH Aachen University, which was sponsored by the National Research Foundation in South Africa and the Federal Ministry of Education and Research (BMBF) in Germany. Since the issue of mineimpacted water is going to be with us for a long time to come, we foresee more such conferences being organized in the future by these well-known higher education and research institutions in collaboration with the SAIMM on a regular basis.

It is my greatest wish that you all enjoy reading the papers in this edition of the Journal, and I hope that you will benefit from some of the ideas presented by the authors.

S. Ndlovu Professor of Metallurgical and Materials Engineering DSI/NRF SARChI Chair: Hydrometallurgy and Sustainable Development School of Chemical and Metallurgical Engineering University of the Witwatersrand, Johannesburg, South Africa

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President’s Corner

When one tugs on a single thing in nature, one finds it attached to the rest of the world

C

hange resistance is the tendency for something to resist change, even when a surprisingly large amount of force is applied. Systemic change resistance is the tendency for a system to reject an attempted change, even though the change is supported over a long period by a substantial fraction of the population. Systems ecologists have increasingly concluded that the conservation of species in isolation from human beings does not address the real systemic issues we are facing. For over 40 years, conservationists have been promoting the need for humans to heed the impact of our activities on the sustainability of natural ecosystem. Despite dire messages from scientists, the system has resisted substantive change.

The theory of change resistance reminds us that, frequently, a culture shift or systemic change process failed because the root cause was not being addressed. The concept of sustainable processing and social responsibility is one such systemic change that is required. Systems ecologists have increasingly warned that the Earth’s ability to sustain the natural ecosystem is at a tipping point, with about 15% of all land globally degraded or severely damaged through human impact. About three out of five people in the world are impacted by damage to the ecosystem. The cost is largely hidden as it is not directly measured or counted, but it is estimated to be equivalent to about 10% of all global wealth.

ESG (environmental, social, and corporate governance) is a top priority for shareholders and investors in mining and metal extraction. Conservation and sustainability messages are not new, however, but it’s become increasingly clear that we cannot protect an elephant without protecting the grass it walks on – both are key components of the ecosystem and cannot be protected in isolation.

Conservationists have tried, and dismally failed. If we attempt to conserve the ecosystem without considering and incorporating human activities into the sustainability plan, the efforts to protect individual species will continue to fall short of the desired goals. Herein lie valuable insights for sustainable and responsible mining and processing. Consider the words of American naturalist John Muir: ‘When we try to pick out anything by itself, we find it hitched to everything else in the Universe’.

Individual professionals will require new skills and perspectives to support corporate ESG targets, and most importantly to truly deliver the systemic changes required to ensure ESG is not just another checkbox exercise in an annual report, so that the principles of sustainability and social responsibility become embedded in how we work. Mining and metallurgical industries can remain viable and deliver sustainable growth only through responsible, ethical, and sustainable mining and processing activities. The SAIMM can play a key role here by creating platforms to advance this crucial agenda beyond the buzzword level. Ultimately, responsible and sustainable processing will be brought about through the professionally influenced behaviour of professionals, such as the members of the SAIMM.

I.J. Geldenhuys President, SAIMM

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Sulphate removal technologies for the treatment of mine-impacted water M. van Rooyen1, P.J. van Staden1, and K.A. du Preez1

Affiliation: 1Mintek, South Africa.

Correspondence to: M. van Rooyen

Email: mouton.michelle@gmail.com

Dates:

Received: 25 Feb. 2021 Published: October 2021

How to cite:

van Rooyen, M., van Staden, P.J., and du Preez, K.A. 2021 Sulphate removal technologies for the treatment of mine-impacted water.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 121, no. 10, pp. 523–530 DOI ID: http://dx.doi.org/10.17159/24119717/1541/2021 ORCID: P.J. Staden https://orcid.org/0000-00018347-3935 This paper was first presented at the Mine-Impacted Water from Waste to Resource Online Conference, 10 and 12, 17 and 19, 3 and 24 November 2020

Synopsis Mine-impacted water, including acid mine drainage (AMD), is a global problem. While precipitation of dissolved metals and neutralization of acidity from mine-impacted water is accomplished relatively easily with lime addition, removal of sulphate to permissible discharge limits is challenging. This paper presents a high-level comparison of four sulphate removal technologies, namely reverse osmosis, ettringite precipitation, barium carbonate addition, and biological sulphate reduction. Primarily operating costs, based on reagent and utility consumptions, are compared. Each process is shown to be subject to a unique set of constraints which might favour one over another for a specific combination of location and AMD composition. Access to and cost of reagents would be a key cost component to any of the processes studied. The total cost calculated for each process also depends on the type of effluents that are allowed to be discharged.

Keywords acid mine water, sulphate removal, reverse osmosis, ettringite, barium carbonate, and biological reduction.

Introduction Acid mine drainage (AMD) arises when pyrite comes into contact with oxygenated water where surfaces have been exposed during mining operations. The pyrite undergoes oxidation in a two-stage process; the first stage produces ferrous sulphate and sulphuric acid, causing the dissolution of metals from surrounding surfaces, while the second stage produces orange-red ferric hydroxide (Johnson and Hallberg, 2005; McCarthy, 2010; van Rooyen and van Staden, 2020).

The acid that forms is partially neutralized by naturally occurring minerals, such as dolomite.

The ferric cations produced can also oxidize additional pyrite and reduce it to ferrous ions and more sulphuric acid (Blodau, 2006; Wolkersdorfer, 2008; van Rooyen and van Staden, 2020). The resultant AMD contains high concentrations of dissolved metals, sulphates, and acidity.

The various technologies that have been developed for AMD treatment can be categorized into physicochemical treatment processes, chemical treatment processes, and biological processes. This paper presents a high-level comparison of four sulphate removal technologies that have received attention in South Africa over the past years. These technologies are reverse osmosis (RO) (physicochemical), the SAVMIN process involving ettringite precipitation (chemical), sulphate removal with barium carbonate addition (chemical), and biological sulphate reduction (biological). Mass balance simulations were completed for each technology and operating costs, based on reagent and utility consumptions, compared. No attempt is made at calculating capital cost but, as a proxy, some comments are given regarding overall process complexity.

The need for sulphate removal AMD typically contains high concentrations of dissolved metals and sulphates, along with low pH values. In South Africa, sulphate concentrations higher than 3000 mg/L in the AMD are common. The South African environmental discharge limits for sulphate-containing water are typically between 250 and 500 mg/L sulphate, depending on the catchment area (van Rooyen, 2016; Neale, 2018). Within the Gauteng region, AMD is currently treated by lime addition to neutralize acid and precipitate the metals.

The sulphate is partially removed, and the gypsum-saturated water, containing sulphate concentrations The Journal of the Southern African Institute of Mining and Metallurgy

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Sulphate removal technologies for the treatment of mine-impacted water of around 1500 mg/L, is discharged into the surrounding water bodies. This process should be applied typically as a precursor to further sulphate removal steps. However, reduction of the sulphate content below that of gypsum-saturated water is more complicated than simple lime treatment and adds to the cost. This issue is aggravated by the environmental challenges associated with the additional waste that is generated.

Several technologies have been proposed for AMD treatment, each producing a different quality of product water. It is essential when choosing an AMD treatment technology that the end use of the product water should be identified first (Mottay and van Staden, 2018). A few typical applications of treated water and the required sulphate levels thereof are given in Table I (Oelofse et al., 2012).

The cost-effectiveness of the process and the waste streams it will generate are also critical factors when choosing a suitable AMD treatment technology. Waste stream generation needs to be minimized to ensure that another potential hazard is not created.

Overview of sulphate removal technologies Approach Mass balance simulations using the IDEAS® simulation platform, supplied by ANDRITZ, were developed for each of the technologies, which enabled a fair comparison between the four processes.

Reverse osmosis The HiPRO® process, implemented at the eMalahleni Water Reclamation Plant in South Africa, is the membrane-based case that was evaluated here. The information from Karakatsanis and Cogho (2010) was used for the preparation of a simplified diagram of the process, shown in Figure 1. The contaminated mine water is first treated with ozone to oxidize iron and manganese, which renders these metals sparingly soluble and causes them to precipitate, to be removed from the suspension in ‘Clarification-1’. This is followed by a reverse osmosis step (‘RO1’) during which the aqueous phase becomes supersaturated with sulphates. However, the addition of antiscalant reportedly keeps the sulphates in solution or, at least, in suspension so that they do not precipitate on the membranes. The brine from ‘RO-1’ is then limed, causing precipitation of the supersaturated ions, followed by ‘Clarification-2’ and ‘RO-2’, and similarly, ‘Clarification-3’ and ‘RO-3’.

The sulphate recoveries for the individual stages are reported to be 70% for RO-1, 65% for RO-2 and 60% for RO-3, resulting in a total recovery of >70% (0.3 × 65%) + (0.3 × 0.35 × 60%) = 9 5.8% for this particular process. The maximum achievable water recoveries from the RO steps are limited by the scaling potential of species such as CaSO4, CaCO3, and SiO2.

During the clarification steps, both gypsum and ‘sludge’ are collected. The former is dewatered before disposal, while the sludge and brine (from RO-3) are disposed of without further treatment, supposedly stored in evaporation ponds. No comment was provided by Karakatsanis and Cogho (2010) regarding alternative dewatering possibilities for the sludge and brine.

Nevertheless, the brine would need to be evaporated either naturally or mechanically, and any solid discard would need to be stored on lined tailings storage facilities.

Reverse osmosis is the only process that is industrially applied for the treatment of acid mine drainage in the South African context. The membrane-based separation would also remove monovalent ions such as Na+ and Cl-. There exist possible processing routes for disposal of the resulting brine, although for the case study considered here no treatment beyond discharge to evaporation ponds has yet been implemented.

Ettringite precipitation Ettringite is a Ca-Al-SO4 mineral with a very low solubility constant, which can be formed by chemical reaction between aqueous CaSO4, such as would occur in the solution phase in equilibrium with precipitated gypsum, and aqueous Al, which can be represented by Al(OH)3. Hence by adding a watersoluble aluminium salt to sulphate-saturated water, the sulphate concentration can easily be reduced to below permissible discharge concentrations.

A process utilizing ettringite precipitation for sulphate removal is described by Nevatalo, van der Meer, and Kerstiens (2014). A more sophisticated ettringite-based process is Mintek’s SAVMIN®, with the distinguishing feature that the aluminium is recovered from the ettringite precipitate for re-use (van Rooyen, 2016). Although significant modifications have been made to the flows heet since its first inception, the concept is most clearly illustrated by the flow diagram provided in Figure 2.

In the ‘Gypsum and metal precipitation’ step, the AMD is first treated with lime to neutralize the acidity (hence generating gypsum) and precipitate the heavy metals as their respective hydroxides. The aqueous overflow contains CaSO4 at a concentration commensurate with the solubility of gypsum. In the ‘Ettringite precipitation’ step, Al(OH)3 is introduced (as a recycle

Table I

cceptable sulphate levels for potential applications of A product water (Oelofse et al., 2012) Product water application

Acceptable sulphate concentration (mg/L)

Coal processing plant General industrial use Discharge to public streams Irrigation Potable use Cooling water in power station

1000 500 500 200 250 20–40

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Figure 1—Block flow diagram of the eMalahleni Water Reclamation Plant, adapted from Hutton et al. (2009) The Journal of the Southern African Institute of Mining and Metallurgy

Sulphate removal technologies for the treatment of mine-impacted water

Figure 2—Block flow diagram of the SAVMIN process (van Rooyen, 2016)

from the ‘Ettringite decomposition’ stage), which incorporates both the calcium and sulphate in solution into solid-phase ettringite. The decant contains only residual Ca(OH)2 alkalinity, which is removed by carbonation with CO2 to precipitate CaCO3 that can be filtered out to yield the treated water product.

The aluminium is recovered from the ettringite by acidification, decomposing the ettringite to Al(OH)3 and gypsum.

The key to the realization of the SAVMIN process has been to achieve the separation of these two solids. There is practically no relative density difference to rely on, the densities being around 2.5 for Al(OH)3 and about 2.3 for gypsum. Therefore, the separation needs to rely on differences in particle shape and size. A multitude of conventional and relatively novel separation apparatus have been tested, with various degrees of success.

Ultimately, hydrocyclones of a suitable design were found to yield the most efficient and most reliable performance. Interestingly it is the slightly higher-density Al(OH)3 that reports to the overflow to be recycled, and the gypsum reports to the underflow to be discarded.

Any gypsum that becomes entrained in the Al(OH)3 recycle stream detracts from the effective aluminium recycle efficiency since the gypsum adds to the sulphate load of the ‘Ettringite precipitation’ step. During a recent piloting campaign, it was possible to restrict the gypsum entrainment in the recycle to <10%.

Logically, any aluminium exiting with the gypsum discard represents a loss that needs to be replenished. This is achieved most effectively by the addition of Al2(SO4)3 to the Ettringite decomposition’ stage since it yields Al(OH)3 that is chemically reactive and does not introduce monovalent elements such as Na+ or Cl- into solution. Indications to date have been that aluminium losses can be restricted to <20%.

A logistical matter to be managed in the SAVMIN process is that all solutions throughout the process are scaling and the formation of gypsum and/or CaCO3 on surfaces is inevitable. This requires sufficient redundancy and duplication of equipment and piping to permit descaling without interruption of the operation.

Note that ettringite precipitation would not remove highly soluble monovalent species such as Na+. However, that does not pose a considerable problem for AMD treatment since Na+ has been found at 100–200 mg/L in AMD which, according to Holmes (1996), poses no health effects although it imparts a faintly salty taste as the concentration approximates 200 mg/L. Precipitation would also not remove highly soluble anions such as Cl-, which accelerates metal corrosion. In the AMD sampled to date, moderate Cl- concentrations around 100 mg/L have been found.

The Journal of the Southern African Institute of Mining and Metallurgy

Barium carbonate addition Based on the information provided by Hlabela, Maree, and Bruinsma (2007) and Mulopo (2015), a simplified diagram of the process is provided in Figure 3. By reducing BaSO4 to BaS with carbon at 1050ºC, while at the same time calcining the CaCO3 that accompanies it, a mixture of BaS and CaO is produced.

When water is added to this mixture, the BaS dissolves, and lime (Ca(OH)2) can be separated from it to be used in the subsequent neutralization step. Carbonation of the BaS solution with CO2 (in the ‘Carbonation 1’ step) strips the sulphur out of solution as H2S, converting the BaS to BaCO3. From the H2S/CO2 gas mixture emanating from the ‘Carbonation-1’ step, elemental sulphur is produced according to the Pipco process, which is discussed by Maree et al. (2005). In essence, this involves burning a portion of the H2S in air to form SO2, which is reacted with the balance of the H2S to yield elemental sulphur (S0) and water.

The BaCO3 thus formed is the key to the process, since it is during the ‘Precipitation’ step that sulphate from the neutralized mine water is captured as highly insoluble BaSO4. Reliance is placed on the extremely low solubility product of BaSO4 to maintain very low concentrations of Ba2+ and SO42- in solution.

The aqueous product of the ‘Precipitation’ step is carbonized with CO2 (in the ‘Carbonation 2’ step) to convert the residual Ca(OH)2 to CaCO3. The mixture of BaSO4 and CaCO3 is separated from the suspension emanating from the ‘Carbonation 2’ step and recycled to the furnace. A small portion of the final water product is recycled for suspending the BaS/CaO solid-phase product of the furnace.

Our calculations suggest that a stoichiometric amount of 0.5 kg carbon (C) is required per cubic metre of typical AMD treated for conducting the chemical reactions in the furnace.

These reactions withdraw oxygen from the BaSO4 and CaCO3 to form CO2. However, another 1.7 kg of carbon per cubic metre is required for reaction with oxygen (introduced in the form of air) to achieve the reaction temperature of 1050ºC. That is based on our simulations, according to which the off-gas from the furnace is used to pre-heat the air supply to about 900ºC to economise the heating requirement.

In the paper by Hlabela, Maree, and Bruinsma (2007), the CO2 emanating from the furnace is shown as being used in the two carbonation steps. However, achieving that in practice seems complex. Since the CO2 generated in the furnace will have been diluted with nitrogen, it will be contaminated with dust and water vapour and will be evolved at about ambient pressure. It would be more feasible in practice to supply the two carbonation steps with pressurized CO2 from a tank, although this would add to the cost and carbon footprint of the process.

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Figure 3—Block flow diagram of the integrated barium process (Mulopo, 2015)

Barium carbonate is highly toxic and catastrophic consequences would result if any unreacted barium carbonate were to pass into the final water product by overdosing the ‘Precipitation’ step with BaCO3. As with the ettringite precipitation processes, precipitation with Ba would not remove monovalent species such as Na+ or Cl-.

Biological sulphate reduction

Biological treatment of mine effluents and mine-impacted water offers a cost-effective and sustainable alternative to conventional treatment technologies. The process has been shown to remove more than 95% of the sulphates, and treated water meets the stringent South African discharge limits for sulphate (Neale et al., 2017). Biological sulphate reduction (BSR) processes employ sulphate-reducing bacteria (SRB) which are present in natural environments such as the sediment of lakes and wetlands, and cattle rumen and subsequent manure. These microbes utilize sulphate as the terminal electron acceptor in anaerobic respiration to produce sulphide. Simultaneously, the sulphide binds with dissolved metals to form stable metal sulphides which are retained in the reactor, reducing concentrations in the treated effluent to trace amounts. There is also a potential for selective recovery of the retained metal sulphides.

SRB also require an electron donor in the form of simple organic substrates, including lactate and acetate (Jamil and Clarke, 2013). These organic substrates are converted to bicarbonate, which increases the pH of the treated water.

The primary reactions are as follows: [1] [2] BSR processes produce significantly less solid waste, with decreased toxicity and increased stability, compared to conventional chemical precipitation methods (Grewar, 2017).

Capital costs for passive BSR are relatively low, and operating costs can be significantly reduced by using inexpensive carbon sources and passive treatment pond designs. Passive BSR processes require little to no energy input and require little maintenance, making them suitable for remote sites post mine closure, as well as for ownerless or legacy mine sites. Active processes, such as THIOPAQ®, have higher operating costs due to the use of costly simple substrates such as hydrogen. These processes are typically employed to produce a high-value product, which offsets the operating costs.

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In South Africa, six BSR plants have been designed and operated at scale: ➤ Mintek’s Biological Sulphate Reduction (BSR) Process, piloted at a local colliery with a pilot study completed in 2019 after 18 months’ operation (Neale, Gericke, and Mülbauer, 2018) ➤ Mintek’s cloSURETM pilot plant is currently being operated at a second mine site, and future activities will include scale-up to a demonstration plant during 2021–2023 ➤ Semi-passive BSR system, demonstrated over 18 months at New Vaal Colliery (van Hille et al., 2016) ➤ Integrated Managed Passive Treatment Process (IMPI), implemented at the currently operating Vryheid Coronation Colliery Passive Water Treatment Plant (still in operation) (Molwantwa et al., 2010) ➤ BioSURE, demonstrated at the East Rand Water Care Company’s (ERWAT) Ancor plant (active process, operated for 15 years) (Rose, 2013) ➤ VitaSOFT Process, demonstrated at VitaOne8’s R&D facilities in Pretoria (active process, currently not operational) (Joubert and Pocock, 2016).

Mintek has been developing cloSURETM, a semi-passive biological treatment process integrating BSR and sulphide oxidation for treatment of mine-impacted water with high sulphate concentrations. The purpose of cloSURE is to treat low volumes (1–4 ML/d) of mine-impacted waters to produce water that is fit for use in irrigated agriculture. Laboratory tests have shown that the two-stage process can produce water of suitable quality for irrigation. The BSR stage has been successfully piloted, and the integrated cloSURE process is currently being piloted at a mine site in Mpumalanga.

Comparison of economics Approach The proponents of the various AMD treatment processes typically do not publish costing information, and the best that can be provided here are indications of where the major cost elements could occur, based on our calculations and simulations. The SAVMIN and BSR processes are exceptions since they are under the control of the developers and hence more accurate cost calculations can be provided. All costs are expressed as South African rands (R).

The attempt to compare the economics of the various processes is further complicated by the difficulty of estimating the capital costs. Only operating costs are provided, with some qualitative comments on likely comparative capital costs.

The Journal of the Southern African Institute of Mining and Metallurgy

Sulphate removal technologies for the treatment of mine-impacted water Neutralization As a basis for the comparison, the AMD collected on the Witwatersrand has been considered as the target water to be treated. It contains 4.7 g/L SO42- in the form of 4.1 g/L Fe2(SO4)3, 0.25 g/L of each of FeSO4 and MgSO4, and lesser quantities of H2SO4 and other metals, to the extent that 3.7 kg Ca(OH)2 per cubic metre of AMD is required to neutralize all acid, precipitate all metals, and raise the pH to 12 (represented by 0.9 g/L Ca(OH)2 in solution). Regardless of the initial elemental AMD composition, lime neutralization would yield a suspension of precipitated metal hydroxides and gypsum in a solution saturated with sulphate, i.e. bearing about 1.5 g/L SO42- (or 2 g/L CaSO4). Assuming a lime cost of R3.0 per kilogram, the cost of neutralization per cubic metre of such an AMD amounts to:

3.7 [kg/m ] x 3.0 [R/kg] = R11/m 3

3

That cost would apply to all processes that would be preceded by neutralization, which include at least the ettringite precipitation, barium carbonate, and BSR processes. During RO, nearly the same cumulative neutralization cost would be incurred over the various precipitation steps. Although some sulphate remains in solution in the ultimate brine product of RO, it should not affect the lime requirement considerably. A significant proportion (67%) of the lime addition in this case is consumed by iron precipitation. The cost comparison is simplified in that only the removal of sulphate from gypsum-saturated water (i.e.

2 g/L CaSO4) needs to be considered as the duty of each process.

The capital cost of neutralization can qualitatively be stated to be ‘low’, with the plant consisting of only a single reactor operating under ambient conditions.

Reverse osmosis RO may generally be associated with high pumping cost due to its historical application in desalination where, for example, an osmotic pressure of around 50 to 60 bar might be encountered to recover 50% fresh water from seawater. This is confirmed by the IMS-Design software provided by Nitto (2019), and from the notes provided by Voutchkov (2008). That would theoretically require a pumping power of 2.2 kWh/m3 which, at an energy price of R1.0 per kilowatt-hour, would amount to a pumping cost of R2.2 per cubic metre. However, in the application of sulphate removal, the IMS-Design software suggests that the osmotic pressure, and therefore the pumping cost, will be about an order of magnitude smaller and will, therefore, be a relatively insignificant contributor to the total operating cost.

Our calculations indicate that a major contributor to the energy cost is ozone generation. According to the information provided by Lenntech (2019), ozone generation requires 16.7 kWh per kg O3. For the application considered here, the stoichiometric requirement for ozone amounts to 0.2 kg/m3, and once again assuming an energy cost of R1.0 per kilowatt-hour, the minimum cost of ozone generation (assuming 100% ozone utilization) amounts to:

0.2 [kg/m3] x 16.7 [kWh/kg] x 1.0 [R/kWh] = R3.3/m3

The process described by Karakatsanis and Cogho. (2010) discards 4.00 m3 brine per 96 m3 water product, assuming 96% clean water recovery as discussed above. A cost needs to be assigned to the treatment or storage of the brine to place the comparison of economics on an equal basis to that of the other processes discussed, which do not produce brines. If it is simply assumed that the brine was to be evaporated (from a liquid at The Journal of the Southern African Institute of Mining and Metallurgy

25ºC to saturated vapour at 100ºC), to crystallize the contained solids for containment, that would involve an energy cost of: 4.0 [m3 brine] x 714 [kWh/m3 brine] x 1.0 [R/kWh] / 96 [m3 water product] = R30 per cubic metre Insufficient information has been available to the authors to estimate other costs such as maintenance labour and materials (which would intuitively be expected to be significant), which can only be provided by those with experience of such an operation.

The capital cost is qualitatively regarded as ‘medium’, requiring multiple membrane separations and clarification steps, all operating at ambient temperature.

Ettringite precipitation Our experience on a 48 m3/d pilot plant operated during 2018/19 has been that the significant cost components are (i) the Ca(OH)2 required for the ettringite precipitation stage and (ii) the acid (H2SO4) needed for the ettringite decomposition stage. With consumptions of respectively 2.3 kg/m3 of Ca(OH)2 and 2.9 kg/m3 of acid (corresponding to the gypsum entrainment and aluminium loss performance indicated earlier) and respective reagent costs of R3.0 and R3.5 per kilogram, the reagent cost amounts to R17 per cubic metre.

SAVMIN is expected to require a relatively small capital expenditure since it operates at ambient pressure and temperature and the plant consists essentially of tankage and piping. Furthermore, since this process does not produce a brine, the disposal and treatment of waste are simplified, consisting essentially of gypsum and metal hydroxide removal.

Barium carbonate addition It is difficult to estimate a cost for this process as no information is yet available on its performance under industrial conditions.

Furthermore, according to the description by Mulopo (2015) of the operation of a 0.5 m3/d pilot plant, a suitable furnace had not yet been identified for conducting the carbon-reduction reaction.

Further uncertainty is introduced by the fact that the recovery of BaCO3 from the ‘Carbonation-1’ step using CO2 does not proceed stoichiometrically. It is not clear what the optimal excess of CO2 and what the associated percentage recovery of BaCO3 would be.

As alluded to above, it would probably be required in practice to add fresh CO2 to the carbonation steps, as opposed to utilizing the CO2 generated in the furnace. No attempt has been made to attach a cost to the CO2 footprint associated with the furnace off-gas for the case where the CO2 from the furnace is not utilized. Assuming a price for CO2 of R0.5 per kilogram and 60% utilization of CO2 in each carbonation step yielding a consumption of 3 kg/m3, the cost of that reagent would amount to R1.5 per cubic metre AMD. Our calculations suggest a carbon consumption (for both chemical reduction and heat generation) of 2.3 kg/m3 and hence, assuming a price for anthracite (bearing 95% carbon) of R1.6 per kilogram, it contributes R3.7 per cubic metre. Presumably, the recovery of BaCO3 is not 100% so that some fresh BaCO3 needs to be added to the process, but the published information does not give any indication of what the amount might be. BaCO3 is a relatively expensive reagent, according to internet information around R9.0 per kilogram.

The stoichiometric requirement for BaCO3 to treat the AMD composition used for this study is 4.7 kg/m3. It follows, therefore, that a 25% loss of CaCO3 would contribute about R11 per cubic metre to the treatment cost.

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Sulphate removal technologies for the treatment of mine-impacted water The requirement for a furnace and the multiple solid-liquid separation steps indicated in the process flow diagram would be expected to render the capital cost higher than that of the SAVMIN process. The capital cost cannot currently be quantified better than by assuming it qualitatively to be ‘high’ compared to that of the other processes discussed here.

2010; Bailey, Gandy, and Jarvis, 2016). With supplementation of an additional carbon source to the bed when necessary, the operational costs will increase depending on the chosen electron donor. For example, the addition of cow manure could increase operating costs to R7.6 per cubic metre, and the use of molasses to R11 per cubic metre.

However, Mintek has developed cloSURE using waste substrate materials, such as cow manure, wood chips, and straw as an electron donor (Neale et al., 2018) to minimize these operating costs. Mintek is currently investigating alternative waste materials for use in the process and an alternative bed design to reduce the costs and challenges associated with organic substrate provision. The research team is also developing a business model for the provision of feedstocks to reduce cloSURE operating costs further.

Biological sulphate reduction The significant challenge for the implementation of BSR processes at scale is the provision of the electron donor to ensure sustainable sulphate reduction. Passive BSR systems typically rely on a packed bed of organic materials, which is intended to degrade over time to provide a sustainable substrate for the microbes (Pulles and Heath, 2009). Often this is supplemented with additional organic material, including ethanol or molasses, to ensure sufficient availability of the electron donor for sulphate reduction to occur. In colder climates and with lower sulphate concentrations in mine-impacted waters, these beds can be operated for 10–15 years with some mechanical intervention; however, in warmer climates, the bed can degrade and impact sulphate reduction efficiency in as little as 8–12 months.

The cost of organic feedstocks to supplement the electron donor in a degrading packed bed can become prohibitively expensive, depending on the level of sulphate in the mine water and the local availability of suitable organic substrates. There seems to be consensus in the industry and the literature (Jamil and Clarke., 2013; van Hille et al., 2016) that the cost of the substrate is the primary reason that passive sulphate reduction technologies have not been readily taken up by the market.

For passive BSR processes, where the degrading packed bed is included in the capital costs, operational costs have been estimated at R4.5 per cubic metre (Gusek and Schneider,

Summary A summary of the important aspects related to each process is given in Table II.

Conclusions Mine-impacted water, including acid mine drainage, is a global problem that requires considerable attention and effort to minimize its destructive impacts. While mere metals precipitation and acid neutralization are generally performed as a primary treatment step, at an operating cost (estimated for the case considered here) of about R11 per cubic metre, additional, more complex processes would be required to achieve a water composition that meets permissible sulphate discharge limits.

This paper presented a high-level comparison of four sulphate removal processes that have relevance, particularly in the South African mining context, namely reverse osmosis, ettringite

Table II

Summary of selected process parameters Parameter

Neutralization with lime

Operating cost

R11/m R33/m R17/m TBD* 3

Reverse osmosis

Ettringite precipitation Barium carbonate addition

3

3

Medium

Potential cost drivers Lime

Energy Lime, H2SO4 CO2, make-up (O3 generation and BaCO3, furnace brine evaporation)

Electron donor (organic substrate)

Process complexity Single unit operation,

including tank(s) and settler

RO membrane units, operating at pressure Multiple tanks and settlers.

Passive ponds

Waste considerations

Gypsum and metal hydroxides. Discharge

on lined TSF

Multiple tanks and settlers

Application: Any (economies of scale) throughput

Furnace, with multiple tanks and settlers

Sulphates <200 mg/L, Sulphates <200 mg/L, monovalents still present. monovalents still present

Brine, requiring evaporation Gypsum and metal and crystallization. hydroxides. Discharge Gypsum and metal on lined TSF hydroxides. Solid waste discharge on lined TSF

Technology Commercialized Commercialized Piloted at 48 m3/d readiness Application: [water] First-stage treatment for

metals removal

High

R4.5–11/m3

Relative capital cost Low

Product water Sulphate concentration Potable water characteristics still too high for permissible discharge.

Low

Biological sulphate reduction Low

Sulphates <200 mg/L, monovalents still present

CO2 and S0 emissions.

H2S gas present.

Gypsum and metal hydroxides / Metal sulphides retained carbonates. Solid waste discharge in the bed. on lined TSF Piloted at 0.5 m3/d, furnace selection to be finalized

IMPI process implemented at field scale, 50 m3

High [SO4] influent when monovalents cannot be tolerated in product

High or low [SO4] influent when monovalents can be tolerated in product.

Any (membrane opex increases linearly)

Any (economies of scale)

Likely limited by furnace capacity

Low/medium

* Process not yet demonstrated, hence accurate operational cost is yet to be determined.

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The Journal of the Southern African Institute of Mining and Metallurgy

Sulphate removal technologies for the treatment of mine-impacted water precipitation, barium carbonate addition, and biological sulphate reduction.

While reverse osmosis is the only process that is industrially applied in South Africa for sulphate removal from acid mine drainage, the generation of brine as a waste justifies consideration of alternative process options for the treatment of AMD. Based on our calculations, the energy cost of ozone generation plus brine evaporation could amount to R33 per cubic metre AMD .

SAVMIN, involving ettringite precipitation, offers one potential alternative, with an operating cost (associated with the consumption of lime and H2SO4) of around R17 per cubic metre for the AMD considered for this study. It would need to be adequately engineered for dealing with the scaling nature of the saturated CaSO4 medium in which the plant will operate, which also applies to any process that treats a saturated calcium sulphate solution. However, capital cost should be low compared with other options.

The addition of barium carbonate for sulphate removal as barium sulphate is another potential precipitation-based process, albeit not yet demonstrated at the scale that SAVMIN has. Notably, a suitable furnace still needs to be identified for the calcination/reduction step. Without a fully defined process, it cannot be costed easily. However, it might suffer from both a high CO2 consumption and a high discharge of CO2 during practical application, and make-up BaCO3 could add to the high cost.

Of the four processes compared in this paper, biological sulphate reduction offers the lowest operating cost, estimated to be between R4.5 and R11 per cubic metre, with the potential for Mintek’s cloSURE technology to reduce the cost of the electron donor further. BSR is a slower process and requires more land than the other technologies and is more suitable for lower throughput volumes. However, its application is well suited to sites post mine closure and legacy sites in remote locations with limited infrastructure, with the potential to re-use the water in irrigated agriculture.

Of the processes considered here, reverse osmosis remains the only process that would also remove monovalent species such as Na+ and Cl-, leaving the process selection dependent on the specific application, or requiring in some instances a combination of processes.

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JOHANNESBURG AND ITS HOLY MINING HERITAGE by

T.R. Stacey and G.J. Heath ABOUT THE BOOK

The City of Johannesburg is probably unique in the world, in that it is a major urban area that was developed on and around a historical goldfield, discovered in 1886. Today the old mine workings run directly beneath the central business district and adjacent areas, as illustrated in the attached photograph showing open stopes exposed in the basement excavation for a building. It is probable that very few residents of the city know that they cross old mine workings, which run at a depth of about 180 m beneath the full length of the M2 motorway, on a daily basis. Although most of the mine entrances were closed when mining ceased, the quality of the closure was often inadequate. Over time, many holes into the old mine workings have appeared on surface, which often present a significant hazard. The increased urbanization that has taken place in South Africa over the past two decades has led to a shortage of land for housing, and the open areas that had been left undeveloped due to the undermining, have seen the growth of informal settlements.

The preparation of this book was prompted by an investigation of old mine openings in the Johannesburg and Central Witwatersrand area by Greg Heath, owing to the potential hazard to residents of the informal settlements. This research, a Council for Geoscience project, was written up as a Master’s dissertation at the University of the Witwatersrand, supervised by Dick Stacey. Although the information is therefore in the public domain, it is not easily accessible by the general public; therefore publication was motivated.

During the project 244 mine openings were located and the degree of hazard assessed. As a result of the evaluation, 80 of the more dangerous openings were subsequently sealed, and this work is also documented in the book. The investigation and sealing of these holes provided a substantial volume of information, which is considered to be of significant historical and scientific value. One of the purposes, then, of this book is to illustrate Johannesburg’s historical mining environment, and to ensure that the valuable recent and historical information that has been collected is preserved in an easily retrievable document.

In addition to details of the investigation and sealing project, additional published information associated with the undermining in Johannesburg, considered to be of historical interest, is included. This material embraces several major development projects across reef outcrops, and cases of subsidence associated with past mining, as well as other examples that are considered to be of interest. Much of this information is based on undermining projects carried out by Dick Stacey and colleagues while employed by SRK Consulting.

A seminal SAIMM Journal paper by Dr F.G. Hill, entitled ‘The stability of the strata overlying the mined-out areas of the central Witwatersrand’, is reproduced in its entirety in an Appendix, acknowledging the major contribution that Dr Hill made to the safety of development works over undermined ground in Johannesburg. Also included are undermining case studies from Dr Tony Brink’s book ‘Engineering Geology of Southern Africa’.

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Are pit lakes an environmentally sustainable closure option for opencast coal mines?

A.C. Johnstone Affiliation: 1GCS (Pty) Ltd, South Africa.

Correspondence to: A.C. Johnstone

Synopsis The aim of the study is to determine if pit lakes are a sustainable coal mine closure option in South African. The water balance, chemistry, limnology, and bacterial population of three selected pit lakes were investigated. The lakes are in the three major coal basins of South Africa and are associated with different lithologies and mining methods. The major factors driving the water balance of the pit lakes are direct rainfall, runoff, inflow from old mine workings, and groundwater infiltration, with the major losses being evaporation or discharge onto surface. The study indicated that pit lakes can be designed as ’terminal sinks’ to provide a sustainable mine closure option. The pit lakes sampled have an alkaline pH, and mostly a sodium/calcium sulphate water with total dissolved solids content of less than 3000 mg/l. The phytoplankton and microbiological data indicates that the pit lakes support aquatic life.

The study shows that correctly designed pit lakes can be an environmentally sustainable closure option for South Africa’s coal mines. A suggested design manual has been developed to assist mine owners and regulators in developing sustainable coal mine pit lakes as a closure option.

Email:

andrewj@gcs-sa.biz

Dates:

Received: 1 Mar. 2021 Revised: 4 Jun. 2021 Accepted: 5 Jun. 2021 Published: October 2021

How to cite:

Johnstone, A.C. 2021 Are pit lakes an environmentally sustainable closure option for opencast coal mines?

Journal of the Southern African Institute of Mining and Metallurgy, vol. 121, no. 10, pp. 531–536 DOI ID: http://dx.doi.org/10.17159/24119717/1551/2021

This paper was first presented at the Mine-Impacted Water from Waste to Resource Online Conference, 10 and 12, 17 and 19, 3 and 24 November 2020

Keywords coal mine. pit lakes, water quality, water balance, mine closure.

Introduction Coal mining started in South Africa in the early 1800s, initially by conventional underground methods, but since the 1950s the majority of the coal production has been from opencast operations. Coal supplies 95% of South Africa’s power needs and will probably continue to do so through the first half of the 21st century. In addition, South African coal production for export and domestic non-power generating purposes is estimated to be 150 Mt/a.

Opencast coal mines generally leave a final void due to the mining method, or insufficient overburden to fill the void left by mining, or as a water management strategy. When the mining operations cease, the void fills with water and forms a lake, which is generally referred to as a ’pit lake’.

The author estimates that there are over 120 pit lakes in the three major coalfields, namely the Witbank, KwaZulu-Natal, and Waterberg coalfields. This study evaluates the environmental sustainability of using pit lakes as a closure option for coal mines in South Africa. The current South African mining and environmental legislation states that in order for a mine to achieve closure, all pit lakes should be backfilled.1 The major factors that determine the environmental sustainability of a pit lake are the water balance and quality. A positive water balance results in discharge from the pit lake onto the surface. A further consideration for environmental sustainability is the chemical and biological nature of the water. Pit lake water quality varies depending on the geology, mining method, and catchment characteristics.

In general, pit lake water quality may not comply with legislated catchment water quality standards.

This investigation considered three pit lakes with the focus on the two major drivers of pit lake sustainability, namely water quality and water balance.

The pit lakes studied were selected on the basis that they are representative of the three major South African coalfields, considering differences in geology and climatic conditions. The factors affecting pit

1

See list of references for applicable legislation

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Are pit lakes an environmentally sustainable closure option for opencast coal mines? lake water balance (and as a result the variation in water level) are groundwater, direct rainfall, and runoff, while the losses from the pit lakes are evaporation, surface discharge, and flow into the surrounding aquifers. The water balance of each of the pit lakes was evaluated to determine the major inputs and losses. The major inputs were identified as groundwater from either the aquifer or backfilled material, and the major loss was evaporation. Pit lake morphology, volume, and surface area are the major design considerations to prevent discharge of lake water into the catchment.

The chemical and biological evolution of the water quality determines the long-term ecological sustainability of pit lakes.

The inorganic chemistry study concentrated on the water quality and vertical stratification. The vertical variation in pH, temperature, dissolved oxygen, and redox potential was measure on a systematic basis in each of the pit lakes. The biotic study investigated the phytoplankton, chlorophyll–a, and the microbiology of each of the lakes.

The waters in the pit lakes are alkaline and have elevated total dissolved solids (mainly calcium sulphate) compared to the ambient surface and groundwater. The pit lakes support life in the form of chlorophyll-a, phytoplankton, microorganisms (bacteria), vegetation, and aquatic life.

A fundamental change in thinking and legislation is required for pit lakes to be accepted as an environmentally sustainable closure option for South African coal mines. This will prevent uncontrolled discharge from opencast mining operations and avoid the expense of ongoing water treatment and associated impacts. Correctly designed pit lakes offer an environmentally sustainable closure option for opencast coal mines in South Africa. Sufficient data was collected in the study to develop a guideline for the design of coal mine pit lakes in the Southern African coalfields. The manual considered the water balance of the pit lakes and the biological and chemical processes that drive the water quality.

Pit lake water balance Pit lake water balances in South Africa are largely controlled by evaporation, as evaporation exceeds precipitation by a factor of 2 to 3. As a result, if the inflow into a pit lake is managed it is unlikely that it will discharge onto surface and into the catchment. The water balances of the pit lakes were calculated based on a generalized mathematical expression (after Gammons et al., 2009):

∆S = (P+ SWin+ GWin) – (E+ (T) + SWout + GWout) where ∆S is change in storage, which is the volume of water in the lake P is the precipitation falling onto the pit lake SWin is the sum of any surface water inputs, which includes runoff and diverted streams GWin is groundwater entering the lake E is the evaporation from the lake T is plant transpiration, which is often negligible SWout is surface water leaving the pit lake, and includes pumpage G Wout is the groundwater leaving the pit lake SWin can be managed by minimizing runoff into the pit lake.

GWin can be minimized (by allowing groundwater levels to rebound) so that evaporation will exceed the sum off all inflows into the pit lake and the lake will act as a water sink.

Pit lake water quality A major consideration is the final pit lake water quality on closure of a mine. This affects the environmental classification and as a result the environmental sustainability of the lake. Conceptual models of pit lake geochemistry are described by external and internal processes, with many of the internal processes being mediated by algae and microbes (Gammons et al., 2009).

Figure 1—Chemical processes in pit lakes (Gammons et al., 2009)

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External processes comprise wallrock runoff and inflow from groundwater, and flow from surface and subsurface mine waste facilities. The quality of the inflows into the pit lake also impacts the evolution of the water quality. Figure 1 summarizes the chemical processes that determine the pit lake water quality.

The evolution of a pit lake at a South African coal mine After mining ceases, water levels in the pit lake rebound to approximately pre-mining groundwater levels. The inflows into the pit lake are rainfall, groundwater, and surface runoff. As water levels rebound the relative contribution of groundwater to the overall pit lake water balance decreases. Rebounding groundwater enters the pit via the dewatered aquifer and from mining activities. The groundwater transports the products of pyrite oxidation, mainly sulphate, iron, and acidity, and dissolves secondary minerals that might have formed in the dewatered aquifer or the backfill material.

Pit lake morphology The pit shape is important from a water balance and quality point of view for the following reasons. ➤ Large surface area compared to mean depth, linear/ elongated pit lake morphology. These pit lakes are more prone to the effects of evaporation and evapoconcentration. Larger surface areas increase losses by evaporation, with lower potential for discharge/overflow, and ensure a groundwater gradient towards the pit lake, to act as a hydrological sink. ➤ Small surface area to mean depth ratio or cone-shaped pit lakes. These pit lakes tend to form an isolated bottom layer, called the monimolimnion, with inferior water quality compared to the rest of the lake. If the chemistry and physical structure of the pit lake are stable, the isolated layer may be beneficial for detaining heavy metals.

Conversely, an isolated layer may not be desirable, as sudden storms events may disturb the chemocline and cause this layer to turn over and mix poor, metal-rich water through the whole water column with deleterious effects on the overall water quality.

Thus, pit lakes with high depth to width ratios are less susceptible to complete lake turnover (Gammons et al., 2009).

Therefore, the relative depth is a good measure for the shape of the pit lake basin (Schultze, 2012). The relative depth is also referred to as the ‘aspect ratio’ and can be described as a comparison of the maximum depth (Zmax; m) of the lake to the lake surface area (Asurface; m2). The relative depth is expressed as a percentage (Castendyk, Eary, and Balistrieri, 2015; Vandenberg, Mccullough, and Castendyk, 2015). The calculation of relative depth is shown in Equation [1]:

Case studies Three case studies, namely the pit lakes at Mafuta, Kriel, and Rooikop, were undertaken. The characteristics of the pit lakes are shown in Tables I and II. Mafutha is single open pit 90 m deep from which 250 000 t of material were removed. The pit is surrounded by an undisturbed aquifer and surface runoff into the pit is largely from the sidewalls. The pit lake has reached equilibrium where groundwater inflow plus direct rainfall equals evaporation, thus resulting in very minor variations in pit lake water levels. The pit lake water at Mafutha is alkaline with a TDS of 1000 mg/l and a sodium chloride type water.

The Kriel site comprises several hydraulically linked open pits left by the extensive opencast mining operations. The pit lakes are hydraulically connected due to the highly permeable backfilled opencast material between the lakes. In this system the major inflows are direct rainfall, runoff, and significantly greater recharge through the overburden spoils. Inflow exceeds evaporation and groundwater outflow, resulting in periodic outflow onto the surface. The pit lake water is alkaline with a TDS of 3500 mg/l and is largely a sodium calcium sulphate water.

The Rooikop pit lake is hydraulically connected to both underground and opencast mine workings. Inflow is from groundwater recharge to the underground workings from the opencast spoils and surface runoff. The pit lake is alkaline with

Figure 2—Morphology of natural lakes compared with pit lakes (Castendyk and Eary, 2009) The Journal of the Southern African Institute of Mining and Metallurgy

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Are pit lakes an environmentally sustainable closure option for opencast coal mines?

Table I

Pit lake physical characteristics

Mafutha Kriel

Coalfield

Rooikop

Waterberg

Witbank /Mpumalanga

Natal

Deep single open pit

Multiple pits hydraulically connected

Single lake

Bulk sample

Opencast rollover method

Opencast and underground

8

Operational to 13 years old

12

Surface area (m2)

26 600

272 530

17 100

Volume (m3)

505 800

797 071

86 681

Dominant inflows

Groundwater, rainfall

Groundwater, rainfall. runoff

Underground mine workings, Runoff from backfill, and natural topography

Dominant outflows

Evaporation

Evaporation

Discharge, evaporation

38.4

3,2

6.8

70

7-10

10

Nature of pit lake Mining method Age (years)

Morphology (%) Depth (m)

Table II

Comparison of pit lake chemistry, phytoplankton, and microbiology measured during the 2016-2017 study period Pit lake

Mafutha

Kriel

Rooikop

pH

8.4

8.4 7.9

Temperature (°C)

27.2 (epilimnion, summer) 18.6 (hypolimnion, summer) 18.6 (whole pit lake, winter)

20.6 (summer) 13.5 (winter)

20 (summer) 15.3 (winter)

Dissolved oxygen (mg/ℓ)

7.53 (epilimnion, summer) 1.3 (hypolimnion, summer) 6.95 (whole pit lake, winter)

R44: 2.03 (summer), 5.7 (winter) R42: 4.9 (summer), 8 (winter)

8.3 (summer) 7.9 (winter)

TDS (mg/l)

1000

3443

1208

Total hardness (as mg/l CaCO3) 186

1210

712

Sodium (mg/l)

301

434

18

Total alkalinity (mg/l)

326

197

118

Sulphate (mg/l)

94

1930

608

Chloride (mg/l)

314

35

2.5

Nitrate-NO3 (mg/l)

9.7

0.24

<0.1

Water type

Na-Cl

Na/Ca-SO4 Ca-SO4

2

10.5

Chlorophyll-a (µg/l)

3.7

Trophic state classification Oligotrophic Mesotrophic to eutrophic Oligotrophic to mesotrophic a

Phytoplankton (dominant phylum,gGenus) Chlorophyta, Ankistrodesmus

a b

Cryptophyta, Cryptomonas and Chlorophyta, Chlorophyta, Ankistrodesmus Chlorella

Microbes (dominating phylum; genera)

Proteobacteria; Acinetobacter, Proteobacteria; Hydrogenophaga, Bacteriodetes; Synechococcus Chlorobaculum, Pseudomonas, Nodularia Flavobacterium, Luteolibacter

Stratification and mixing

• Strong thermal stratification (October to March/ April) • Turnover (late autumn/ winter)

• Weak thermal stratification (October to March/April) • Turnover (late autumn/ winter)

• Weak thermal stratification (October toMarch/April) • Turnover (late autumn/ winter)

Classificationb

Slightly alkaline, low TDS Holomictic, monomictic

Slightly alkaline, high TDS Holomictic, monomictic

Circum-neutral, low TDS Holomictic, monomictic

Trophic state classification according to de Lange et al. (2018). Only September data for Rooikop and Kleinfontein.

Classification of pit lakes according to Eary (1999) and Boehrer and Schultze (2008)

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Are pit lakes an environmentally sustainable closure option for opencast coal mines? a total dissolved solids content of 1200 mg/l and the water is calcium sulphate type.

Boehrer, B. and Schultze, M. 2008. Stratification of lakes. Reviews of Geophysics,

Are pit lakes an environmentally sustainable closure option for South African coal mines?

Castendyk, D.N., Eary, L.E., and Balistrieri, L.A. 2015. Modelling and management of

The use of pit lakes in the South African coal mining industry as an environmentally sustainable closure option must consider the following aspects, which should be incorporated into the closure design.

Water balance It is critical that the sum of the inflows into the pit lake is less than evaporation to prevent major changes in pit lake water levels. Positive pit lake water balances result in discharge onto surface. Negative water balances result in the pit lake becoming a hydraulic sink. As a result, a carefully designed pit lake where the inflows are equal to or less than the outflows will result in a stable pit lake that will not discharge onto surface. The morphology is a critical factor in determining the environmental sustainability of a pit lake.

Water quality The pit lakes studied in this investigation are all alkaline with variable dissolved solids contents and which support phytoplankton, microbes, vegetation, and aquatic life. Although the water quality may not comply with the catchment water quality objectives, the pit lakes are environmentally sustainable.

The pit lakes were also seen to support a host of other aquatic life such as birds, fish, amphibians and mammals.

Conclusion The surface area of a pit lake is vitally important. A large surface area maximizes evaporation, which directly affects the water balance. In addition, surface runoff should be controlled to avoid excess flow into the pit lake during storm events, which may lead to a temporary positive water balance and uncontrolled discharge into the catchment. Should the pit lake be suitably designed, it forms a water sink and prevents uncontrolled discharge from the mining operations.

The water quality in the pit lakes investigated is alkaline with evaluated dissolved solid content, but is able to support chlorophyll-a, phytoplankton, and microbes. Although the water quality may not comply with catchment water quality standards, it supports a stable ecosystem.

Correctly designed pit lakes are a sustainable closure option for South African coal mines. It is recommended that the current South African legislation be reviewed in order to accommodate pit lakes as a sustainable closure option. However, each pit lake is unique and as a result must be evaluated on a case-by-case basis.

Details of the study can be found in the South African Water Research Commission report TT 797/1/19 (Johnstone, Kennedy, and Mpetle, 2019).

References Blanchette, M.L. and Lund, M.A. 2016. Pit lakes are a global legacy of mining: An integrated approach to achieving sustainable ecosystem and value for communities. Science Direct, vol. 23 (December).

vol. 46, RG2005. doi:10.1029/2006RG000210

pit lake water chemistry 1. Theory. Applied Geochemistry, vol. 57. pp. 267–288.

Castendyk, D.N. and Eary, L.E. 2009. The nature and global distribution of pit lakes.

Mine Pit Lakes Characteristics, Predictive Modelling and Sustainability. Society for Mining, Metallurgy and Exploration, Littleton, CO. Doyle, G.A. and Runnells, D.D. 1997. Physical limnology of existing mine pit lakes.

Mining Engineering, vol. 49. pp. 76–80.

De Lange, W.J., Genthe, B., Hill, L., and Oberholser, P.J. 2018. Towards a rapid assessment protocol for identifying pit lakes worthy of restoration. Journal of Environmental Management, vol. 206. pp. 949–961.

Eary, E. 1999. Geochemical and equilibrium trends in mine pit lakes. Applied Geochemistry, vol. 14. pp. 963–987.

Gammons, C.H., Harris, L.N., Castro, J.M., Cott, P.A., and Hanna, B.W. 2009. Creating lakes from open pit mines: Processes and considerations - with emphasis on northern environments. Canadian Technical Report of Fisheries and Aquatic Sciences 2826. https://digitalcommons.mtech.edu/cgi/viewcontent. cgi?article=1001&context=geol_engr Huisamen, A. and Wolkersdorfer, C. 2015. Modeling the hydrogeochemical evolution of mine water in a decommissioned opencast coal mine. International Journal of Coal Geology, vol. 164. http://dx.doi.org/10.1016/j.coal.2016.05.006 International Council on Mining & Metals. 2019. Integrated Mine Closure Good Practice Guide. http://www.icmm.com/en-gb/guidance/environmental-stewardship/ integrated-mine-closure-2019 Johnstone, A.C., Kennedy, L., and Mpetle, M. 2019. An investigation to determine if South African coal mine pit lakes are a viable closure option. Water Research Commission publication TT 797/19, volumes 1 & 2. Pretoria.

Schultze, M. 2012 The filling and remediation of pit lakes in former open cast lignite mines. PhD thesis, Technical University of Carolo Wilhemina, Braunschweig.

Snyman, C.P. 1998. Coal. The Mineral Resources of South Africa. 6th edn. Wilson, M.G.C. and Anhaeusser, C.R. (eds). Handbook 16, Council for Geoscience, Pretoria.

South Africa. Not dated. The South African National Mine Closure Strategy (in process).

South Africa. 2015. The Financial Provisioning Regulations, 2015.(Government / notice R1147, Government Gazette 39425, 20 November 2015.

South Africa. 2002. Mineral & Petroleum Resources Development Act, 28 of 2002 and the Regulations 56(e).

South Africa. 1998a. National Environmental Management Act, 107 of 1998 and the EIA Regulations. (Government No. 19519, Notice No. 1540. Commencement date: 29 January 1999 [Proc. No. 8, Gazette No. 19703]).

South Africa. 1998b. National Water Act, 36 of 1998.

South Africa. 1996. Constitution of the Republic of South Africa.

Vandenberg, J., McCullough, C., and Castendyk, D. 2015. Key issues in mine closure planning related to pit lakes. Proceedings of the 10th International Conference on Acid Rock Drainage and IMWA Annual Conference, Santiago, Chile. https:// www.imwa.info/docs/imwa_2015/IMWA2015_Vandenberg_156.pdf Wetzel, R.G. 2001. Limnology: Lake and River Ecosystems. 3rd edn. Academic Press, San Diego, CA. 1006 pp.

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SAIMM ONLINE SHORT COURSE

Are pit lakes an environmentally sustainable closure option for opencast coal mines?

Environmental Constraints in Blasting

ONLINE SHORT COURSE 25 FEBRUARY 2021

CPD Points 0.1 ECSA CPD Points for every 1 hour webinar attended

BACKGROUND Blasting is necessary for the recovery of ore, minerals or stone in our underground and surface mines and also in the development of key infrastructure in civil blasting. However, blasting can cause noise and vibration, which can have an impact upon neighbouring premises. Proper control of blasting practices is therefore necessary to ensure both the safety of employees and the protection of the community from adverse effects. Ground vibration, noise, airblast and flyrock, usually results in complaints from the public following blasting operations. More often than not, excess noise and airblast are the cause of the complaint even though the complainant may have in fact claimed that ground vibrations are the problem. A considerable amount of research into blasting vibrations and consequential damage has been done. In most cases, a competent operator can reasonably predict the level of airblast and ground vibration. However the generation and transmission of airblast and ground vibration is affected by many factors. The challenge forward is not to just apply a limit rather to promote a better understanding of the key leverage points to better manage the impact of blasting and the unique local constraints impacting on the mine operation. We will use a real life example to understand these constraints, how to manage them and the impact on the way we do blast design in our short course.

SAIMM have responded to our member feedback and are excited to be offering short-courses to develop critical technical skills across the complete mining value chain. We are excited to offer our second course in this series focused on Environmental Constraints in Blasting.

This short-course is being offered from 17h00 until 19h30 so that you can take advantage of your time in lockdown and do some learning.

CPD points towards your Continual Professional Development Requirements. Please add the usual CPD wording.

PROGRAMME • • • • • •

17h00-17h10 17h10-18h00 18h00-18h10 18h10-19h00 19h00-19h30 19h30

Welcome and Introduction Session 1 Comfort break Session 2 Open Discussion Closing

PRESENTER Simon Tose is an expert in Blast Design from AEL and will be sharing his expertise on this subject.

Simon Tose is an established industry recognized consultant, registered professional engineer and leader in technology, mining, explosives and blasting science. Holding a BSC Hons Mining Engineering degree, management and explosive qualifications. Extensive experience in mining methods, education and management of projects and blast investigations. He has written papers, articles and presented at a number of International and Local conferences. Current board member of the IOQSA & ISEE and regular contributions to the SAIMM & AusIMM. With a strength in project design he leads the Blast Consult team with a strong passion for the development of environmental and blast monitoring, measurement and investigation, consulting, management and financial analysis for AEL.

FOR FURTHER INFORMATION, CONTACT: Camielah Jardine, OCTOBER 2021 ▶ 536 Head of Conferencing

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The effect of decarburization on the fatigue life of overhead line hardware J. Calitz1, S. Kok2, and D. Delport3

Affiliation: 1 Tshwane University of Technology, South Africa and Eskom - Research, Testing and Development, South Africa.

2 University of Pretoria, South Africa.

3 Tshwane University of Technology, South Africa.

Correspondence to: J. Calitz

Email:

calitzj@eskom.co.za

Dates:

Received: 4 Feb. 2020 Revised: 25 Feb. 2021 Accepted: 13 Sep. 2021 Published: October 2021

Synopsis Altering the microstructure in order to improve the tensile properties of bow shackles resulted in inconsistency in the fatigue performance. This raises the question whether the inconsistency in fatigue life can be attributed to microstructural changes along the profile of the shackle or to decarburization at the surface.

Bow shackles forged from 080M40 (EN8) material were subjected to different heat treatments in order to alter the microstructure. The shackles were subjected to five different fatigue load cases, which represented typical loads experienced at termination points for an overhead power line with a span length of 400 m, with changes in conductor type, configuration, wind, and ice loading.

Although the change in microstructure does improve both the tensile and fatigue performance, we found that the depth of the decarburization layer has a greater effect on the high cycle fatigue life of bow shackles than the non-homogeneous microstructure.

Keywords bow shackle, decarburization, fatigue, hardness, microstructure.

How to cite:

Calitz, J., Kok, S., and Delport, D. 2021 The effect of decarburization on the fatigue life of overhead line hardware.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 121, no. 10, pp. 537–542 DOI ID: http://dx.doi.org/10.17159/24119717/1109/2021 ORCID: J. Calitz https://orcid.org/0000-00022719-9406 S. Kok https://orcid.org/0000-00019453-2806 D. Delport https://orcid.org/0000-00031963-556X This paper was first presented at the Mine-Impacted Water from Waste to Resource Online Conference, 10 and 12, 17 and 19, 3 and 24 November 2020

Introduction In order to reduce the cost of an overhead power line, larger overhead conductor diameters and longer line span lengths (distance between two structures) are used, hence reducing the number of structures used for a specific line length (Calitz et al., 2005).

The prime function of overhead line hardware is to connect the phase conductor to the insulators and the insulators to the structure (tower) (Eskom, 2014). The design philosophy of reducing the number of structures results in an increase in the mechanical loading on termination points at structures, especially at termination (strain) structures, as a result of the increase in conductor mass (Calitz et al., 2005). Consequently, the line hardware strength classes have to increase.

Quenching and tempering heat treatment methods are typically used to increase the mechanical strength of line hardware, which often result in a nonhomogeneous microstructure along the profile, particularly of forged components with a complex geometry, due to the difference in cooling rates for the different cross-sectional areas (Wieser, 1980).

Therefore, when bow shackle failures occurred as a result of fatigue damage on certain span lengths (sections) on the 765 kV electrical network in particular, it raised the question what impact poor control during the heat treatment of bow shackles has on their fatigue performance.

This study investigates the influence of microstructure and decarburization on the fatigue behaviour of forged line hardware, specifically bow shackles.

Literature survey Gildersleeve (1991) investigated the relationship between decarburization and fatigue strength of a manganese-molybdenum through-hardening steel (605M36). Rotating bending tests were used to determine the fatigue behaviour. The specimens had decarburized layers up to 1 mm in depth.

The results showed that the fatigue limit was mostly independent of the depth of decarburization.

Gildersleeve also examined surface carbon concentration and found the fatigue limit to be linearly dependent upon the carbon concentration at the surface.

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The effect of decarburization on the fatigue life of overhead line hardware Adamaszek and Broz (2001) investigated the effects of decarburization on hardness changes in carbon steels caused by high-temperature surface oxidation. They found that the decarburization resulted from annealing, and caused the grains near the surface to grow. The authors explain that during the decarburization process oxygen penetrates the surface scale through cavities, pores, and cracks. This oxygen reacts with the different elements in the metal, causing decarburization. The extent of decarburization is greater for metals with a higher Fe content. The authors also found hardness changes due to the decarburized layers, resulting in lower fatigue resistance.

Hankins and Becker (1932) tested forged specimens of four types of steel of different hardnesses which included both low- and medium-carbon steels, under cantilever-type rotating bending. The hourglass-shaped specimens consisted of both asforged surface finish specimens, from which the flash had been trimmed before testing, and machined and polished specimens.

A metallurgical analysis of the test specimens showed that the hardness of the decarburized surface was lower than that of the interior material. The specimens with the largest difference between surface and interior hardness displayed the largest difference in endurance limit between the polished and as-forged surface conditions.

The authors state that decarburization is the main cause of reduced fatigue life for specimens in the as-forged surface condition.

Experimental methods

Heat treatment of test samples The normalizing heat treatment was conducted in a protective (controlled) atmosphere using propane gas to prevent scaling of the component surface. The shackles were placed in baskets and loaded into the preheated furnace at 820°C. Subsequently, the temperature was increased to 920°C within 75 minutes.

The temperature was maintained for 2½ hours, after which the components were removed and air cooled to room temperature while still in the baskets. Figure 2 depicts the heat treatment equipment used.

In order to alter the microstructure the shackles were subjected to two different hardening and tempering heat treatments, namely ‘AR’ (as received) and ‘HTM’ (heat treatment modification). ➤ Austenitization heat treatment: Following the normalizing heat treatment, the steps were:

• A R: Reheated to 880°C for approximately 40 minutes, followed by quenching in oil which was maintained at a temperature range between 60°C and 80°C.

• H TM: Reheated to 880°C for approximately 1 hour, followed by quenching in running water. ➤ Tempering heat treatment: The last stage of the heat treatment process was tempering, which was conducted at:

• A R: 515°C for 4 hours followed by oil quenching

• H TM: 520°C for 4 hours followed by quenching in water.

Testing

Manufacturing of test samples Bow shackles were hammer-forged from 35 mm diameter 080M40 (EN8) material, using a closed die system making use of a Banning pneumatic hammer forged press as depicted in Figure 1. The 300 mm long billets were induction heated between 1100°C and 1250°C and the forging temperature was between 950°C and 1250°C. The holes in the eye section were punched immediately after the component was forged.

After heat treatment the shackles were galvanized using a centrifuge galvanizing method. The bath had a zinc content of 99.7% and the temperature of the molten zinc was maintained at 450°C. The shackles were removed before solidification of the zinc could occur on the component surfaces and placed in a centrifuge, then spun for several seconds to remove excess zinc from the surface. Thereafter, the components were transferred to a quench tank where they were cooled to allow handling.

The high cycle fatigue testing was conducted on a MTS Landmark servohydraulic test machine. As testing was not conducted on standard fatigue specimens, but on components (the bow shackles), the test requirements of ASTM E466–96 (ASTM, 2002) were used as a guideline to develop a force-controlled constant amplitude axial fatigue test procedure in MTS TestSuite Multipurpose Elite (mpe) software, which was used for setting the test parameters and recording the test data.

As depicted in Figure 3, a back-to-back shackle arrangement was used to simulate installation practices. The tongue fittings measuring 20 mm in thickness and a bolt hole of 20 mm was used to connect the back-to-back shackle arrangement to the MTS test machine grips.

The bolts used were not the standard 8.8 grade bolts that are supplied with the 210 kN shackles, but were machined from martensitic stainless steel, which provided improved mechanical

Figure 1—Induction furnace and hammer forge press used to manufacture the bow shackles

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Figure 2—Furnace arrangement used for the austenitization heat treatment and quenching process

Figure 3—Shackle fatigue testing set-up

properties compared to the 8.8 grade material, in order to minimize the deformation associated when subjecting the bolt to bending loads when the shackle is loaded in tension.

The selection of which hardware strength class to use for a hardware assembly is determined by the ultimate tensile strength of the conductor attached to it; thus hardware rated at 210 kN can accommodate different conductor types and configurations.

Therefore the mechanical loading introduced to the hardware assembly is directly linked to the tension within the conductor, which is affected by several factors such as conductor type, temperature, span length, wind, and ice loading.

In-house developed software titled Tower Loader SANS v2.3d, based on SANS 10280 (SANS, 2013), was utilized to calculate the different loading conditions for a 210 kN rated hardware assembly. The loading conditions were subdivided into fatigue load cases which are summarized in Table I. The amplitude (lower and upper limits) is expressed in load (kN) based on the percentage of the ultimate tensile strength of the shackles, which was taken as 210 kN. The Journal of the Southern African Institute of Mining and Metallurgy

After each fatigue test, the test specimen was sectioned approximately 20 mm from the fracture end in order to measure the hardness of the material. In all cases, the leg section failed.

In addition, specimens were removed from the eye and crown sections. The specimens were cut using a Labotom-15 manual cut-off machine with cutting/cooling fluid. Both transverse and longitudinal specimens were removed.

Table I

Fatigue load cases Fatigue

load case 1 2 3 4 5

Lower limit

(kN) 37.8 52.5 71.4 100.8 121.8

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Upper limit

(% of UTS)

(kN)

(% of UTS)

18 25 34 48 58

52.5 71.4 100.8 121.8 147.0

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The effect of decarburization on the fatigue life of overhead line hardware The metallographic preparation included the hot mounting of specimens in an ATM Opal 460 hot mounting press at 180°C and 250 bar pressure using thermosetting bakelite resin in order to improve the handling of the specimens.

To remove the surface damage caused by cutting, the mounted specimens were ground with rotating discs of silicon carbide paper. The grinding procedure involved several stages, using a finer grit paper with each subsequent stage in order to remove the scratches from the previous coarser paper. The last stage was conducted with a 1200 µm grit paper, followed by two stages of polishing with diamond suspension of 3 and 1 µm grit to produce a smooth surface finish.

After cleaning with running water and acetone, the specimens were etched with 5% Nital, a solution of nitric acid and methanol, in order to reveal the microstructure of the steel. The specimens were immediately washed with methanol and dried after etching.

A computer-controlled inverted metallurgical Zeiss Axio microscope with ZEN 2 core imaging software was used to conduct the metallographic examinations and to capture digital images of the microstructure, at different magnifications.

Vickers hardness measurements were taken across the prepared surfaces of both transverse and longitudinal specimens,

utilizing a semi-automated Emcotest DuraScan 70 hardness machine with Ecos work flow software. Measurements were taken at 1 mm intervals by applying a load of 10 kgf.

Results Figure 4 depicts the number of cycles to failure versus the mean applied load for the ‘as received’ shackles.

Figure 5 depicts the microstructures of the metallographic specimens removed from the leg section of shackle specimens A and B in the vicinity where the fatigue fracture occurred. The specimen with the higher ferrite content, namely shackle B, underwent a higher number of cycles to failure. As depicted in Figure 5, shackle B also had a decarburization depth 38% less than that of shackle A. As depicted in Figure 6, shackle B had an overall lower hardness than shackle A. However, shackle B revealed better fatigue performance than shackle A, as depicted in Figure 4. This can be attributed to shackle A having a larger decarburization depth and greater hardness difference between the surface and the parent (inner) material.

The altered microstructure of the HTM shackles that was obtained by the modified heat treatment resulted in an

Figure 4—Mean load versus number of cycles to failure plot for ’as received’ shackles

Figure 5—Difference in decarburization depth for ’as received’ shackles

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Figure 6—Hardness profiles for ‘as received’ shackles A and B

improvement in the fatigue performance of the bow shackles for loading cases 1 and 2, as illustrated in Figure 7. No improvement in fatigue performance was gained at higher loadings. A previous study (Calitz, Kok, and Delport, 2019) concluded that the curved shape of the leg contributes towards premature fatigue failure at higher loads The leg section is not only subjected to tensile stress when the shackle is under tension, but also to a large degree of bending stress as a consequence of its shape. The increase in bending stress with an increase in loading results in localized overstressing of the material, promoting fatigue crack initiation and propagation.

Figure 8 depicts the hardness profiles and micrographs taken at the leg section of shackle C. Shackle C represents the overall condition of the HTM shackles. The decarburization depth was less prominent for the HTM shackles and the hardness difference between surface and parent material was minimal compared to shackle A (Figure 6).

Conclusion The fatigue performance of a bow shackle can be improved by altering its microstructure. However, this improvement The Journal of the Southern African Institute of Mining and Metallurgy

by metallurgical means is limited to lower load cases, as the curved shape of the leg contributes towards premature fatigue failure. The leg section is not only subjected to tensile stress when the shackle is under tension, but also to a large degree of bending stress as a consequence of its shape. The increase in bending stress with an increase in loading results in localized overstressing of the material, promoting fatigue crack initiation and propagation.

In addition, the fatigue performance of the shackles at lower stress levels was significantly influenced by the depth and hardness of the decarburization layer. A decrease in hardness and an increase in the depth of the decarburization layer resulted in poor fatigue performance, even with an increase in the through thickness hardness of the leg section.

This difference in hardness is one explanation for the decreased life in the high cycle fatigue (HCF) region for forged shackles, since higher hardness (higher strength) is a desired property for long fatigue life.

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Figure 7—Mean load versus number of cycles to failure plot for HTM shackles

Figure 8—Hardness profiles for HTM shackle C

is the main cause of reduced fatigue life and not the nonhomogenous microstructure.

Acknowledgements The lead author would like to acknowledge Eskom Holdings SOC Limited for their commitment to providing him with the opportunity to further his engineering skills and knowledge.

Through Eskom’s vision in the Eskom Power Plant Engineering Institute (EPPEI) programme he is able to pursue his doctoral degree in engineering.

References Adamaszek, K. and Broz, P. 2001. Decarburization and hardness changes in carbon steels caused by high temperature surface oxidation in ambient air. Diffusion and Defect Data: Defect and Diffusion Forum, vol. 194. pp. 1701–1706.

ASTM. 2002. E 466-96. Standard practice for conducting force controlled constant amplitude axial fatigue tests of metallic materials. ASTM International, West Conshohocken, PA.

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Calitz, J., Haridass, B., Jacobs, B., Retief, J., and Plessis, P.D. 2005. Line hardware.

The Planning, Design and Construction of Overhead Power Lines. Bisnath, S., Britten, A. C., Cretchley, D. H., Muftic, D., Pillay, T., and Vajeth, R. (eds).

Crown Publications, Johannesburg. pp. 441–482 Calitz, J., Kok, S., and Delport, D. 2019. The effect of geometry on the fatigue life of overhead line hardware. Journal of Failure Analysis and Prevention, vol. 19, no. 5. pp. 1401–1406.

Eskom. 2014. Specification for suspension and strain assemblies and for hardware for transmission lines. Sandton, South Africa.

Gildersleeve, M. 1991. Relationship between decarburization and fatigue strength of through hardened and carburizing steels. Material Science and Technology, vol.

7. pp. 307–310.

Hankins, G. and Becker, M. 1932. The fatigue resistance of unmachined forged steels.

Journal of the Iron and Steel Institute, vol. 126. pp. 205–236.

SANS. 2013. SANS 10280-1. Overhead power lines for conditions prevailing in South Africa. Part 1: Safety. SABS Standards Division, Pretoria: pp. 30, 33.

Wieser, P.F. 1980. Supplement 11: Hardenabilty and heat treatment. Steel castings Handbook (5th edn). Steel Founders' Society of America. u The Journal of the Southern African Institute of Mining and Metallurgy

Microstructure, microhardness, and tensile properties of hotrolled Al6061/TiB2/CeO2 hybrid composites Affiliation: 1 Department of Mechanical Engineering PES Institute of Technology, India.

2 Department of Industrial Engineering & Management, Ramaiah Institute of Technology, India.

3 Department of Mechanical Engineering, National Institute of Technology Karnataka, Surathkal, India.

Correspondence to: P.G. Koppad D. Sethuram

Email:

praveennath2007@gmail.com sethuramD@pes.edu

S. Iyengar1, D. Sethuram1, R. Shobha2, and P.G. Koppad3

Synopsis TiB2 and CeO2 particle-reinforced Al6061 hybrid composites were manufactured using stir casting and hot rolling techniques. The base alloy and composites were hot-rolled at 500ºC and a 50% reduction was achieved through 12 passes. The effect of varying TiB2 and CeO2 particle additions on the microstructure and mechanical properties of the Al6061 matrix was studied. Scanning electron microscopy showed uniform dispersion of both the reinforcements, with good interfacial bonding. Microhardness and tensile properties like yield and tensile strength were found to be higher for hybrid composite with 2.5% TiB2 and 2.5% CeO2 compared to Al6061 alloy and other hybrid composites. The increased tensile strength is attributed to good dispersion and interfacial bonding between the particles and Al6061 matrix. Fracture analysis using a scanning electron microscope revealed ductile fracture for the Al6061 alloy and mixed characteristics of ductile-brittle fracture for hybrid composites.

Keywords aluminium matrix composites; hot rolling; microstructure; mechanical properties; fracture analysis.

Dates:

Received: 5 Mar. 2020 Revised: 24 Aug. 2021 Accepted: 16 Sep. 2021 Published: October 2021

How to cite:

Iyengar, S., Sethuram, D., Shobha, R., and Koppad, P.G. 2021 Microstructure, microhardness, and tensile properties of hotrolled Al6061/TiB2/CeO2 hybrid composites.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 121, no. 10, pp. 543–548 DOI ID: http://dx.doi.org/10.17159/24119717/1560/2021 ORCID: P.G. Koppad https://orcid.org/0000-00034872-2638 This paper was first presented at the Mine-Impacted Water from Waste to Resource Online Conference, 10 and 12, 17 and 19, 3 and 24 November 2020

Introduction The development of lightweight metal matrix composites with multiple reinforcements has been extensively studied owing to their numerous advantages over single reinforcement metal matrix composites. Among all lightweight metals, aluminium and its alloys have been considered as potential candidate materials for the matrix phase (Surappa, 2003; Bodunrin, Alaneme, and Chown, 2015).

Aluminium and its alloys have low density, good corrosion resistance, and moderate to high strength.

In particular Al6061 alloy, which contains major alloying elements such as magnesium and silicon, is known for good formability and moderate strength (Bray, 1990). At present this alloy is used for construction and transport applications such as rail coaches, bridge railings, couplings, valves, welded structures, and helicopter rotor skins. The base Al6061 alloy can be reinforced with different filler materials to improve its properties. The most commonly used reinforcement materials are ceramic and carbon-based particles/fibres (Ram, Koppad, and Kashyap, 2014; Rajesh, Auradi, and Kori, 2016).

Although the early studies focused on the development of single reinforcement-based composites, the past couple of decades have seen an appreciable amount of work on hybrid composites (Sharma, 2000; Ramesh et al., 2010; Boppana et al., 2020; Ghazanlou, Eghbali, and Petrov, 2021).

The newly developed hybrid composites not only have good mechanical properties, but also exhibit good physical and tribological properties (Vellingiri, 2019; Mallikarjuna et al., 2020). They are capable of meeting both load-bearing and wear resistance requirements for many applications. Mummoorthi, Rajkumar, and Kumar (2019) reported high tensile strength and elongation of 421 MPa and 11.2% respectively for Al6061 hybrid composites with 5% Fe2O3 and 6% B4C particles. The wear rate tends to decrease with increased hybrid reinforcement content. Prakash, Sivasankaran, and Sasikumar (2015) found that the Brinell hardness of friction-stir processed Al6061 hybrid composites increased up to 6% Al2O3 content, decreasing thereafter.

The performance of hybrid composites depends largely on the appropriate selection of reinforcement combinations. The addition of soft phases like graphite, graphene, or carbon nanotubes reduces the coefficient of friction as well as the wear rate of hybrid composites. Premnath et al. (2014) studied the hardness and wear behaviour of Al6061/Al2O3/Gr hybrid composites developed using stir casting. With

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Microstructure, microhardness, and tensile properties of hot-rolled Al6061/TiB2 increased Al2O3 particle content from 5% to 15% the hardness increased marginally. Wear testing was conducted as per the central composite design method, and ANOVA was used to obtain significant wear parameters. Load was identified as the dominant factor affecting the wear rate of the composites.

As mentioned earlier, the enhancement of mechanical and tribological properties is highly dependent on the hybrid reinforcement combination, but the effect of the processing route also needs to be considered. An appropriate processing route helps in achieving good bonding between matrix and reinforcement and uniform dispersion of the reinforcement in the metallic matrix. Much research has employed thermomechanical processing to obtain good quality and enhanced properties of hybrid composites. In particular, hot rolling is used to shape cast composites into sheet or plate form, as well as improve the properties. Hot rolling is generally carried out above the recrystallization temperature and is capable of refining the microstructure significantly; it also helps in obtaining good bonding and uniform dispersion of reinforcements. Kumar et al. (2018) employed hot rolling to develop Al6061/ZrB2 in-situ composites. The rolled composites showed a significant reduction in the grain size and improvement in the tensile properties compared to as-cast composites. The composite with 10% ZrB2 content showed the highest tensile and yield strength values of 324 and 310 MPa. Xu et al. (2012) studied the effect of rolling reduction on the tensile properties of Al6061/Al18B4O33 composites. Increased rolling reduction from 30% to 70%, improved the tensile properties. Additionally, Al18B4O33 whiskers tended to orient themselves along the rolling direction. Nie et al. (2017) investigated the mechanical properties of hot-rolled Al-TiB2/TiC hybrid composites. Rolling was carried out at 300ºC with different rolling reductions of 20% to 90%. After rolling, the microstructures revealed improved particle distribution with significant grain refinement. Similarly, the hardness and strength increased with increased rolling reduction.

Exploiting the advantages of hot rolling, this work focused on the development of Al6061/TiB2/CeO2 hybrid composites using stir casting and hot rolling. The addition of multiple reinforcements in the Al6061 matrix can cause clustering in cast conditions. In order to break reinforcement clusters, hot rolling was used for grain refinement and to improve the mechanical properties. The effect of varying reinforcement contents on the mechanical properties of Al6061 alloy is presented.

then poured into permanent moulds of cast iron. After casting, all the samples were subjected to hot rolling using a two-high rolling mill (Buhler Group, Germany). Prior to hot rolling, all the samples were heated to 515ºC for about one hour. After heating, hot rolling was carried out to achieve about 50% rolling reduction in a 12-pass rolling schedule. The Al6061 and hybrid composites with their designations are presented in Table II.

Characterization and testing Hot-rolled samples were cut by electrical discharge machining (EDM). Samples for scanning electron microscopy (SEM), microhardness, and tensile tests were machined using EDM as per ASTM standards (ASTM E8/E8M, 2016). For microscopic analysis, the samples were polished using grit papers followed by 0.25 µm diamond paste. To reveal microstructures such as grain boundaries, all the samples were etched using Keller’s reagent (190 ml distilled water, 5 ml HNO, 3 ml HCl, 2 ml HF). Scanning electron microscopy with energy dispersive X-ray analysis (EDX) (Hitachi, model number SU3500N) was used for the analysis of the hot-rolled microstructure, reinforcement dispersion, interfacial bonding, and composition. The micrographs were taken in backscattered (BSE) and secondary electron (SE) modes for analysis of microstructural features. Microhardness tests were conducted using a Vickers tester (Shimadzu microhardness tester) employing a load of 500 g. For each sample, about five indentations were made and the average value recorded.

Tensile test samples were prepared as per ASTM E8/E8M and tests were conducted using a Universal testing machine. A schematic diagram of the tensile test sample is provided in Figure 1. For each sample, about three tests were conducted and the average value recorded. Fracture analysis after tensile tests was conducted on fractured samples of Al6061 alloy and its hybrid composites using SEM.

Results and discussion Microstructure analysis The microstructures of the hot-rolled Al6061 alloy and hybrid composites are shown in Figure 2. Figure 2a shows the Al6061 alloy, displaying elongation and orientation of aluminium grains along the rolling direction. Generally, in the as-cast condition Al6061 alloy had a dendritic microstructure, but during hot rolling the grains with the dendritic structure tended to break down, forming elongated grains with fine equiaxed grains.

Experimental Table I

Materials and method The Al6061/TiB2 composites with 2.5 and 5 wt% TiB2 particle content were manufactured by in-situ reaction between Al-3%B and Al-10%Ti master alloys. The chemical composition of the Al6061 alloy used as matrix material is presented in Table I. The detailed procedure for the synthesis of Al6061/TiB2 composite is described in Ramesh, Pramod, and Keshavamurthy (2011).

For these in-situ composites, CeO2 particles at 2.5 and 5 wt% were added externally. The in-situ composite melt was held at 800ºC, followed by gradual addition of preheated CeO2 particles, and stirred using a mechanical stirrer rotating at 400 r/min.

Stirring was continued for about 15 minutes to ensure proper mixing and good dispersions of CeO2 particles. After stirring, hexachloroethane tablets were used to degas the molten metal to avoid casting-related defects. The degassed molten metal was

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Chemical composition of the Al6061 alloy Element

Al

Mg

Si

Cu

Mn

Fe

Weight %

Balance

1.08

0.63

0.32

0.52

0.17

Table II

omposition of hybrid composites and sample C designations No. 1 2 3 4

Sample composition

Designation

Al6061 alloy AA1 Al6061 + 2.5%CeO2 + 2.5%TiB2 HB1 Al6061 + 2.5%CeO2 + 5.0%TiB2 HB2 Al6061 + 5.0%CeO2 + 2.5%TiB2 HB3The Journal of the Southern African Institute of Mining and Metallurgy

Microstructure, microhardness, and tensile properties of hot-rolled Al6061/TiB2

Figure 1—Schematic of a tensile test sample

Figure 2b shows the micrograph of HB1, which contained a fairly uniform dispersion of TiB2 and CeO2 particles. This is the main advantage of hot rolling because it breaks up inhomogeneous dispersions formed during casting by fragmenting them into individual particles and forming good bonding due to the application of compressive forces (El-Sabbagh et al., 2013).

Figure 2c shows HB3, which had an elongated microstructure with many fine equiaxed grains between the elongated grains.

Figure 2d shows HB3 with fairly uniform dispersion of both TiB2 and CeO2 particles. The reinforcements were located at the grain boundaries and few were within the grains. Generally, the as-cast composites showed inhomogeneous dispersion of reinforcements and large particle-free zones. The clusters of reinforcement were very few in the hot-rolled composite, as seen in Figure 2d. This suggests that due to the application of compressive forces most of the reinforcement clusters were broken up. In addition to reinforcement declustering, the hot

rolling redistributed and oriented the grains in the rolling direction. The particle-free zones were quite narrow, showing good distribution of both reinforcement particles. The absence of defects suggests that the compressive forces applied during hot rolling were sufficient to eliminate them. The compressive forces, along with elevated temperature, helped to close the porosity and improve the interfacial bonding between reinforcements and the Al6061 matrix. The clean interfaces, without any processing defects, showed that good bonding was achieved between the constituents of hybrid composites. This is important because the formation of any reaction products at the interface can have detrimental effects on the mechanical properties. Figure 3a shows that for the HB1 hybrid composite there was good interfacial bonding between the two reinforcements and the Al6061 matrix.

The good bonding can be inferred from the fact that there were no defects at the reinforcement/matrix interface.

In general, the cast aluminium composites suffer problems with molten aluminium wetting the reinforcements, especially the ceramic particulates. The unwetted particles trap gases, which leads to the formation of gas voids at the interface, and numerous gas voids may form a network of interconnected voids. In addition, there is a high possibility of clustering of reinforcements due to gas adsorption by the unwetted particles.

Alternatively, detrimental products may form at the interface, which could change the composition of the matrix and affect the final properties of the composites (Hashim, Looney, and Hashmi, 2001; Gawdzinska, Chybowski, and Przetakiewicz, 2016). In the present case, the interface was continuous with no interfacial products, gas voids, or crack formation due to the coefficient of thermal expansion mismatch. Also, neither TiB2 nor CeO2 particles reacted with the Al6061 matrix, which could have led to the formation of brittle reaction products at the interface. The absence of any kind of defects or interfacial products indicates that the bonding between reinforcements and Al6061 matrix was good. The EDX peaks, as shown in Figure 3b, indicate that most of the elements detected correspond to constituents of the hybrid

Figure 2—SEM micrographs of hot-rolled specimens. (a) AA1 taken in BS mode, (b) HB1 taken in SE mode, (c) HB3 showing grain structure, taken in BS mode, and (d) HB3 showing reinforcement dispersion, taken in SE mode The Journal of the Southern African Institute of Mining and Metallurgy

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Figure 3—Hot-rolled HB1 (Al6061+2.5%CeO2+2.5TiB2 hybrid composite). (a) SEM-SE micrograph showing good interfacial bonding between reinforcements and Al6061 matrix, and (b) EDX analysis of the complete area

composite. The presence of Mg and Si in the pattern corresponds to the major alloying elements of the Al6061 matrix. On the other hand, Ti, B, Ce and O correspond to both the reinforcements CeO2 and TiB2 particles. From the EDX analysis it was quite clear that no detrimental interfacial products formed during casting or hot rolling and the interfaces were clean and continuous.

Microhardness

The microhardness values of hot-rolled Al6061 alloy and its hybrid composites were studied to ascertain the effect of multiple reinforcements. Sample AA1, with no reinforcement, had a microhardness value of 57±2 VHN. With the addition of 2.5% CeO2 and 2.5% TiB2, sample HB1 showed a microhardness value of 83±3 VHN – compared to AA1, about 46% increment in microhardness value was recorded for sample HB1. The other hybrid composites, HB2 and HB3, had microhardness values of 66±2 and 71±4 VHN respectively, representing increments of 16% and 25% compared to AA1. Addition of CeO2 and TiB2 particles resulted in a significant enhancement in microhardness of the Al6061 matrix. However, a higher reinforcement content led to decreased microhardness in hybrid composites HB2 and HB3. Overall higher microhardness values for hybrid composites indicate that both hot rolling and reinforcements increased the resistance to deformation. The increment in microhardness is attributed to densification of the microstructure due to hot rolling and the addition of hard reinforcing phases like CeO2 and TiB2 particles.

There is a very high probability of pore formation after casting of hybrid composites. These pores adversely affect the mechanical properties of both Al6061 alloy and hybrid composites, whereas hot rolling tends to improve the properties by eliminating pores. The plastic deformation caused by compressive forces at high temperature was the main reason for the reduction of pores in these samples (Guo et al., 2017).

The other main factor contributing to enhanced hardness was

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resistance to plastic deformation due to the presence of hard reinforcing phases. The uniformly dispersed CeO2 and TiB2 particles tended to resist plastic deformation caused by the indenter of the Vickers hardness tester. The resistance to plastic deformation was due to inhibition of dislocation movement by CeO2 and TiB2 particles (Shin and Bae, 2015). Further grain refinement caused by the hot rolling tends to increase the grain boundary volume. This inhibited the movement of dislocations, thereby contributing to improved microhardness values.

Tensile properties The tensile strength, yield strength (0.2% proof stress), and elongation values of hot-rolled Al6061 alloy and its hybrid composites were studied. Sample AA1 showed yield and tensile strength values of 78±5 MPa and 130±4 MPa respectively. The hybrid composite HB1, which contained 2.5% CeO2 and 2.5% TiB2, showed yield and tensile strength values of 124±4 and 208±3 MPa respectively. The increment is quite significant compared to the unreinforced AA1 sample.

Increments of 60% and 59% in yield and tensile strength values for HB1 demonstrated the synergistic effect of hot rolling and strengthening by CeO2 and TiB2 particles. The HB2 sample showed yield and tensile strength values of 115±7 MPa and 183±6 MPa, while HB3 showed 104±5 MPa and 169±7MPa respectively. All hybrid composites displayed significant enhancement in strength. However, the best reinforcement combination for obtaining high yield and tensile strength was that for HB1, constituting 2.5% CeO2 and 2.5% TiB2 particles.

The strength increase of hybrid composites may be due to the following factors. i. T he hot rolling used for densification of the microstructure not only improves density and closes the casting-related pores, but also helps in uniform dispersion of the reinforcements.

Furthermore, the compressive forces during rolling improve The Journal of the Southern African Institute of Mining and Metallurgy

Microstructure, microhardness, and tensile properties of hot-rolled Al6061/TiB2 the bonding between both the reinforcements and Al6061 matrix. The uniformly dispersed CeO2 and TiB2 particles with good bonding are capable of bearing the load from the matrix because the good bonding helps in the efficient transfer of load from the Al6061 matrix to both CeO2 and TiB2 particles. ii. The difference between the coefficients of thermal expansion of CeO2 (approx. 10.7×10-6/K), TiB2 (approx. 6× 10-6/K) and Al6061 (approx. 23.4×10-6/K) is quite large.

Due to this, dislocations form around the reinforcements.

Uniform dispersion of reinforcements helps in dispering the dislocations uniformly in the matrix. The dislocations formed due to thermal mismatch become entangled with the dislocations already present (Poza and Llorca., 1999), thus increasing the dislocation density, which is another factor contributing to strength enhancement.

As well as strength values, the elongation of Al6061 alloy and its hybrid composites was studied. For Al6061 alloy the elongation was 17.2%. In the case of the hybrid composites, HB1, HB2, and HB3, the elongation values were 9.2%, 13.9%, and 10.1% respectively. The relatively high elongation value for HB2 is due mainly to the high proportion of in-situ formed TiB2 particles (5%). This agrees with the findings of Han, Liu, and Bian (2002) on TiB2 particle-reinforced Al-Si alloy composites, who reported increased ductility as the TiB2 particle proportions increased from 1% to 7%. The grain refinement and improved interfacial bonding of reinforcements with the matrix from hot rolling were the main reasons for the increment in elongation values for hybrid composites.

Fracture analysis Figure 4 shows the SEM micrographs of tensile fractured samples of Al6061 alloy and the hybrid composites. Unreinforced AA1(Figure 4a) underwent ductile-type fracture. The fracture surface had large numbers of dimples, indicating plastic deformation. Figures 4 b, c, and d show the fracture surfaces of hybrid composites HB1, HB2, and HB3. As seen in Figure 4b the fracture surface of HB1 again had a large number of dimples

due to plastic deformation of the Al6061 matrix. However, some regions in the fracture surface of hybrid composite HB1 had no dimples at all, but displayed cleavage facets that formed due to the addition of hard CeO2 and TiB2 particles to the Al6061 matrix. Such cleavage regions showed some brittle fracture, and these hybrid composites exhibited mixed features of ductilebrittle fracture (Manikandan et al., 2020). As observed in Figure 4c, the fracture surface of hybrid composite HB2 was covered with numerous dimples, indicating that the sample had undergone plastic deformation, which is consistent with the highest elongation value of 13.9%. The presence of multiple reinforcements and application of hot rolling resulted in the fine equiaxed grains in the hybrid composites. In addition, the dispersion of reinforcements was quite uniform and bonding with the matrix was good, which is why the dimples were present all over the fracture surfaces. Hybrid composite HB3 had a similar fracture surface to HB1, indicating failure due to the combined effect of ductile-brittle fracture (Figure 4d). The dimple-covered fractured surfaces as seen in Figures 4b to 4d indicate that the hybrid composites were not adversely affected by the addition of the reinforcements and were able to display sufficient ductility (Xie et al., 2020).

Conclusions The effect of hot rolling and addition of multiple reinforcements in the form of CeO2 and TiB2 particles on microstructure, microhardness and tensile properties of Al6061 alloy were evaluated. i.

S EM analysis showed uniform dispersion of CeO2 and TiB2 particles in the Al6061 matrix. The improved bonding between matrix and reinforcements is attributed to hot rolling. ii. T he microhardness of the hybrid composites was significantly better than that of unreinforced Al6061 alloy. The hybrid composite with 2.5% CeO2 and 2.5% TiB2 had the best microhardness, at 83±3 VHN.

Figure 4—SEM-SE micrographs showing tensile fracture surfaces of (a) AA1, (b) HB1, (c) HB2, and (d) HB3 The Journal of the Southern African Institute of Mining and Metallurgy

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Microstructure, microhardness, and tensile properties of hot-rolled Al6061/TiB2 iii. Tensile testing revealed significant increases in yield and tensile strength of hybrid composites with the addition of CeO2 and TiB2 particles. The optimum reinforcement combination for attaining strength was 2.5% CeO2 and 2.5% TiB2 particles. iv. Fracture analysis conducted using electron microscopy revealed that Al6061 alloy underwent ductile fracture, while the hybrid composites exhibited mixed ductile-brittle fracture.

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Nie, J., Wang, F., Li, Y., Liu, Y., Liu, X., and Zhao, Y. 2017. Microstructure and mechanical properties of Al-TiB2/TiC in situ composites improved via hot

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The Journal of the Southern African Institute of Mining and Metallurgy

Resistance spot welding of a thin 0.7 mm EN10130: DC04 material onto a thicker 2.4 mm 817M40 engineering steel Affiliation: 1 Department of Materials Science and Metallurgical Engineering, University of Pretoria, South Africa.

2 Engineering Department, Faculty of MIT, Malawi University of Science and Technology, Malaw.

3 Graduate school of Technology Management, University of Pretoria of Pretoria, South Africa.

K.A. Annan1, R.C. Nkhoma2, and S. Ngomane3

Synopsis

Email:

The effects of welding current, electrode force, and welding time in a resistance spot weld were studied to investigate the effectiveness of welded joints between a thin EN10130: DC04 material and a thicker 817M40 part, through analysis of the microstructural and mechanical properties. All welded specimens were subjected to tensile testing at room temperature (25°C) and sub-zero temperature (–46°C) to test the strength of the welded joints. No full button failure was observed at either room temperature or sub-zero temperature after optimization of the weldng parameters. The fusion zone was observed to consist mainly of martensitic phase, due to rapid quenching, while the HAZ was composed of clusters of martensite in a ferrite and bainite matrix. The base 817M40 metal remained fully ferritic after welding.

The hardness was found to increase with increasing welding current. An increase in nugget size, indicating good fusion of the weld, was observed with an increase in the welding current.

Dates:

Keywords microstructure, resistance spot weld, hardness.

Correspondence to: K.A. Annan

kofi.annan@up.ac.za

Received: 13 Apr. 2021 Revised: 12 Sep. 2021 Accepted: 22 Sep. 2021 Published: October 2021

How to cite:

Annan, K.A., Nkhoma, R.C., and Ngomane, S. 2021 Resistance spot welding of a thin 0.7 mm- EN10130: DC04 material onto a thicker 2.4 mm 817M40 engineering steel.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 121, no. 10, pp. 549–556 DOI ID: http://dx.doi.org/10.17159/24119717/1597/2021 ORCID: K.A. Annan https://orcid.org/0000-00029623-8106 R.C. Nkhoma https://orcid.org/0002-6407-432X This paper was first presented at the Mine-Impacted Water from Waste to Resource Online Conference, 10 and 12, 17 and 19, 3 and 24 November 2020

Introduction 817M40 steel, also known as EN24T, is a Ni-Cr-Mo high hardenability, high tensile strength, and high wear resistance steel. EN24T can be heat treated to obtain a wide range of improved mechanical properties (Ramazani et al., 2015; Khan et al., 2008; Choi et al., 2011). This grade is very popular and widely used for many high-strength applications where a good combination of strength and impact properties is essential in large components (Ramazani et al., 2015; Khan et al., 2008). EN10130:DC04 is, however, a non-alloy steel suitable for applications requiring high ductility such as cold forming of components with complex profiles at high deformation speed (Khan et al., 2008).

Welding of 817M40 steel sheet in the hardened and tempered condition should be carried out with great caution, as the mechanical properties can be altered significantly through the welding process (Ramazani et al., 2015; Choi et al., 2011; Al-Mukhtar and Doos, 2013; Ghazanfari and Naderi, 2013).

The most widely used method for joining two different materials such as EN24T and EN1030:DC04 is resistance spot welding (RSW), due to the fact that different configurations can be obtained with this method (Al-Mukhtar and Doos, 2013; Ghazanfari and Naderi, 2013; Wan, Wang, and Zhang, 2014; Pouranvari, 2011). Factors that influence the RSW process include the current passing through the workpiece, the time for which the current flows through the workpiece, the electrode pressure, and contact area between the electrode tip and the workpiece (Wan, Wang, and Zhang, 2014; Pouranvari, 2011; Singh et al., 2019). The most widely investigated RSW parameters that need to be controlled to achieve a good weld joint include the welding current, time, and electrode force (Ghazanfari and Naderi, 2013). The welding current has been found to be an influential parameter in determining the strength of the weld joint, since an increase in welding current leads to an increased heat flow, which influences the microstructural changes (Ghazanfari and Naderi, 2013; Raut and AchwaL, 2014).

The welding current provides the necessary heat for melting and therefore has a direct bearing on the microstructural changes (Khan et al., 2008; Wan, Wang, and Zhang, 2014; Pouranvari, 2011; Singh et al., 2019; Raut and AchwaL, 2014; Ali, Khan, and Moeed, 2015). It has been found that a higher welding current increases the strength of the weld nuggets due to the increased weld nugget area of the steel. The different weld regions whose microstructure impacts the mechanical properties of the welded

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Resistance spot welding of a thin 0.7 mm EN10130: DC04 material joints include the base metal (BM), the heat-affected zone (HAZ), and the fusion zone (FZ) (Pouranvari, 2011). The microstructure of the FZ or nugget diameter, which is determined by the heat input, also affects the mechanical behaviour of the welded joint.

RSW is one of the electric welding methods that rely on heat, time, and pressure and is widely used in the automotive industry. Determining the settings for heat, time, and pressure is not an easy task, as highlighted by Nayaka et al. (2012 and Nasir and Khan (2016). It is more of an art than science, and is done through experimental test work. Set quality checks such as (shear) tensile tests, penetration, nugget visual inspection, and dimensional checks are used to determine ‘pass’ or ‘fail’ criteria (Singh et al., 2019; Liu et al, 2019; Pouranvari, 2017; Pouranvari, Sobhani, and Goodarzi, 2018).

The environment in which a material is applied also influences the properties of the material, including the strength as well as the microstructure (Ali, Khan, and Moeed, 2015, Pouranvari, 2017; Lee and Chang, 2020). It is, therefore, important to characterize the microstructure evolution during welding and its influence on the final mechanical properties of the steels. It has been reported that the failure mode depends primarily on the size of the fusion zone, and failure tends to occur under the pull-out mode as the fusion zone increases in size (Zhao, Zhang, and Lai, 2018; Sun, Stephens, and Khaleel, 2007). In the current study we investigated the effects of welding current, welding time, and electrode force on the tensile shear strength of resistance spot welded joints at different temperatures.

determined following a series of tests conducted to optimize the welding conditions. The values were selected so as to ensure a minimum nugget diameter of 4√t, where t is the sheet thickness, after welding so that industrial welding conditions were maintained as far as possible (Sun, Stephens, and Khaleel, 2007). Four sets of RSW tests were done at one weld condition in order to evaluate the effect of the parameters on the weld.

Table II presents the parameters used in the welding process through a variation process to ensure a total of 24 tests in a round-robin situation. The holding rate (cooling rate) was kept ~ at a maximum of approximately 1°C/s. Each test was conducted three times to ensure repeatability.

Figure 1 presents a schematic diagram of the welding set-up used.

The samples for testing at sub-zero temperature were placed in dry ice in a vacuum container for 2 hours before being removed for tensile testing. The temperature of the samples was ascertained to be at least below –50°C at the time they were taken from the vacuum container so that the temperature at the testing time did not rise above –46°C. The tensile tests were carried out on an MTS Criterion 45 machine at room temperature (25°C) and sub-zero temperature (–46°C) at a crosshead velocity of 2 mm/min and the interfacial fractured surfaces of the samples examined under a metallurgical stereo microscope. Metallographic

Materials and experimental procedure A 2.4 mm thick 817 M40 steel was used as a base metal and 0.7 mm EN10130: DC04 steel as thin material. The chemical compositions of the steels are presented in Table I. The surfaces of the metals were cleaned with 1.0 M HCl solution and a metal brush to remove any contamination before welding was done.

RSW was performed using a base metal of 2.4 mm thickness and a 0.7 mm thin material. The RSW tests were carried out ~ using 45° truncated cone Cr-Zr- Cu containing electrodes with a 6 mm face diameter from Resistance Welding Manufacturers' Association (RWMA). The bottom electrode (tip) was shaped to follow the contour of the inside diameter of the cylinder for proper current flow and distribution. The parameters used were

Figure 1—A schematic diagram of the welding set-up

Table I

As-received composition of 81740 and DC04 steels used in the investigation (wt%) Steel 817M40 D C04

C

Si

P

S

Al

Ni

Mn

Mo

Cr

0.38 0.242

0.31 0.007

0.01 0.0144

0.01 0.0082

0.03 0.04

1.52 0.028

0.54 0.22

0.32 –

1.24 –

Table II

Welding parameters used in the study Specimen no.

1 2 3 4

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Welding current (kA)

Welding time (cycles)

Electrode force (kN)

6 8 10 12

6 8 10 12

4.5 5.0 5.5 6.0

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Resistance spot welding of a thin 0.7 mm EN10130: DC04 material samples from the pulled weld nuggets were cut with a precision saw, mounted, and ground with a Struers Tetrapol-25 grinding machine using silicon carbide (SiC) paper sizes of 240, 400, 600, 800, and 1200 followed by polishing to a mirror-like finish using diamond suspension down to 1 μm for 3 minutes.

All metallographic samples were prepared in accordance with ASTM 407 standard (ASTM, 2015). The mechanically prepared metallographic samples were etched in 2% Nital and examined under an Olympus BX51M microscope. The fractured surfaces were analysed using a JEOL SEM at working distance of 15 mm and 20 kV to establish the nature of the fracture following the tensile test. Hardness testing was carried out over the fusion zone, HAZ, and the BM using a Future Tech hardness machine.

A diamond-type indenter was used for measurement, with a load of 2 kgf and dwelling time of 10 seconds. The hardness measurement was repeated with three patterns and the average value calculated.

Results and discussion Examination of interfacial fractured surfaces Figure 2 shows images of the samples pulled at different temperatures, illustrating the nature of fracture that occurred. As seen, no full button failure or full pullout was observed in any samples as only partial pullout was recorded. Samples tested at both room temperature and –46°C underwent partial interfacial fracture, as illustrated in Figure 2.

Figures 3 and 4 show the nugget morphology and the macroscopic morphology of the cross-section of the welded joints of samples welded using different conditions and pulled at room temperature and sub-zero temperature. The morphologies show cracks, but no shrinkage hole in the fusion zone for different welding currents used (Ramazani et al., 2015; Nayaka et al., 2012). All cross-section morphologies showed the welding characteristics zones of interest FZ, HAZ, and BM. It was also

Figure 2—Images of the samples pulled at (a) room temperature (25°C) and (b) –46°C

Figure 3—Interfacial fractured surfaces of samples welded with a welding current of (a) 6 kA (b) 8 kA, (c) 10 kA, and (d). 12 kA, and pulled at 25°C The Journal of the Southern African Institute of Mining and Metallurgy

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Figure 4—Interfacial fractured surface of samples welded with a welding current of (a) 6 kA, (b) 8 kA, (c) 10 kA, and (d) 12 kA, and pulled at –46°

Figure 5—Microstructures of the welded joints showing the (a) BM, (b) FZ, and (c) HAZ of samples welded with a current of 10 kA

observed that the nugget diameter after pulling increased with increasing welding current in samples pulled at room temperature as well as those pulled at sub-zero temperature.

An increasing depth of penetration into the parent material was observed as the welding current increased from 6 kA to 12 kA, with values increasing from 1.2 mm to 2 mm, showing good fusion in all the welds. A good level of fusion and penetration was observed in the samples welded with a current of 10 kA. The samples pulled at both 25°C and –46°C all underwent ductile fracture, as shown in the SEM micrographs in Figure 5.

Microstructural observations From the typical cross-section morphologies of the welded joints, three zones of interest namely the FZ, HAZ, and BM, were identified for microstructural examination. The microstructures

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of these zones are shown in Figure 5. The temperature of the BM remains below the transformation temperature of the steel and no transformation took place, as depicted by the microstructure in Figure 5a (Ramazani et al., 2015). The FZ experienced a high temperature during welding and a fast cooling rate thereafter, leading to phase transformation of austenite to martensite (Ramazani et al., 2015; Nayaka et al., 2012; Pouranvari, Sobhani, and Goodarzi, 2018). The martensite has a columnar texture with solidified needle-like structures as shown in Figure 5b. Figure 5c shows the microstructure of the HAZ following rapid post-weld cooling which resulted in the transformation of the austenitized structure to undissolved ferrite and bainite clusters (Lee and Chang, 2020; Zhao, Zhang, and Lai, 2018).

The microstructure of the HAZ is made up of martensite clusters formed by decomposition of C-rich austenite regions in the ferrite The Journal of the Southern African Institute of Mining and Metallurgy

Resistance spot welding of a thin 0.7 mm EN10130: DC04 material matrix. The microstructure of the HAZ changed appreciably with increasing welding current, with a greater volume of martensite and bainite and less ferrite phase formed at higher currents.

The fractured surfaces of the pulled samples were examined using SEM in order to determine the nature of the fracture.

The samples pulled at 25°C and –46°C both underwent ductile fracture, as shown in the SEM micrographs in Figure 6.

Hardness measurements Figure 7 shows the measured hardness values as a function of the welding conditions for samples tensile-tested at room temperature and sub-zero (–46°C). The hardness of all the samples tested at sub-zero temperature is higher than for those tested at room temperature. This increase in hardness can be attributed to the fresh martensite that might have transformed in the sub-zero tensile test samples (Nayaka et al., 2012; Sherepenko et al., 2019; Rao et al., 2017). The effect of the electrode force on hardness (Figure 7b) is lower than that of welding current, (compare with Figure 7a). The increase in martensitic content following fast cooling after welding resulted in increased hardness, as shown in the hardness profiles in Figures 7a and 7c, which also show the effect of welding current and time on the hardness (Sherepenko and Jüttner, 2018; Sherepenko et al., 2019). In Figure 7a, the decrease in hardness with increasing welding current is observed in samples pulled at room temperature and at sub-zero temperature (Rao et al, 2017; Mohamadizadeh, Biro, and Worswick, 2020). A similar profile of softening is also observed with increasing welding time, as shown in Figure 7c.

Tensile properties The influence of welding current, electrode force, and welding time on the nugget diameter, energy absorbed before fracture, as well as the maximum load experienced before fracture are presented in Figure 8 for samples pulled at room temperature and at sub-zero temperature (–46°C). The tests at different temperatures were carried out to evaluate the effect of the environment on the welded joint strength.

An increase in nugget size was recorded with increasing welding current. The maximum load to fracture and failure energy also varied with welding current. Figure 8b indicates that the maximum nugget size was obtained at an electrode force of 5.5 kN. At electrode forces lower than 5.5 kN, expulsions were

recorded due to the decreased heat input, leading to a nugget size that did not provide effective contact. The results also shows that the nugget size and failure performance of the weld joint decrease with an increase in welding time beyond 10 cycles as a result of welding defects (Rao et al., 2017). The samples with increased nugget diameters also exhibited the highest pulling strength, as pullout fracture was observed in these samples.

The decrease in maximum load and fracture energy for samples pulled at sub-zero temperature indicates increased brittleness in these samples (Ramazani et al., 2015; Nayaka et al., 2012; Rao et al., 2017). The brittleness is also confirmed by the increased hardness values for the samples pulled at sub-zero temperatures.

This indicates that the temperature environment within which the welded component is used influences the mechanical properties, which is bound to affect the performance in such environments.

A penetration depth of 81% was observed as the recorded penetration of 2.00 mm in the 2.44 mm parent material. A low penetration rate is associated with a small nugget size, while penetration of 75% and above is reported to indicate improved weld strength quality Ramazani et al., 2015; Choi et al., 2011).

A combination of the parameters led to an increase in the nugget size, showing that each of the parameters has an influence, and only through a careful combination of the required parameters can optimal strength be obtained. In this work the optimum results were obtained at a welding current of 10 kA, electrode force of 5.5 kN, and welding time of 10 cycles.

Conclusions The effects of welding current, welding time, and electrode force on the quality of weld joints between a 2.4 mm EN24T steel and a 0.75 mm EN10130: DC04 steel were investigated. The following conclusions are drawn. ➤ Increased welding current led to an increase in hardness, and a reduction in the fracture energy and maximum load in samples tensile tested at both room temperature and sub-zero temperatures. ➤ Increases in welding current, welding time, and electrode force resulted in an increased nugget size. The increase in nugget size led to an increase in the penetration and strength of the joint. ➤ The welding current is the most significant welding parameter influencing the quality of the weld joint and the type of microstructures obtained

Figure 6—SEM micrographs of the fractured surfaces after pulling at (a) 25°C and (b) –46°C The Journal of the Southern African Institute of Mining and Metallurgy

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Figure 7—Plots of hardness measured after pulling at 25°C and –46°C as a function of (a) welding current, (b) electrode force, and (c) welding time

➤ No full button failure was observed in either the samples pulled at room temperature or those pulled at –46°C after optimization of the welding conditions. ➤ The optimal welding conditons were a welding current of 10 KA, electrode force of 5.5 kN, and welding time of 10 cycles.

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Correlations of geotechnical monitoring data in open pit slope back-analysis – A mine case study A.F. Silva1, J.M.G. Sotomayor2, and V.F.N. Torres2 Affiliation: 1Vale, Parauapebas, Brazil.

2 Vale Institute of Technology, Ouro Preto, Brazil.

Correspondence to: J.M.G. Sotomayor

Email:

juan.sotomayor@itv.org

Dates:

Received: 12 May 2021 Revised: 31 Aug. 2021 Accepted: 6 Sep. 2021 Published: October 2021

How to cite:

Silva, A.F., Sotomayor, J.M.G., Torres, V.F.N. 2021 Correlations of geotechnical monitoring data in open pit slope back analysis – A mine case study.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 121, no. 10, pp. 557–564 DOI ID: http://dx.doi.org/10.17159/24119717/1618/2021 ORCID: A.F. Silva https://orcid.org/0000-00031369-0103 J.M.G. Sotomayor https://orcid.org/0000-00031523-1138 V.F.N. Torres https://orcid.org/0000-00024262-0916 This paper was first presented at the Mine-Impacted Water from Waste to Resource Online Conference, 10 and 12, 17 and 19, 3 and 24 November 2020

Synopsis Geotechnical monitoring plays an important role in the detection of operational safety issues in the slopes of open pits. Currently, monitoring companies offer several solutions involving robust technologies that boast highly reliable data and the ability to control risky conditions. The monitoring data must be processed and analysed so as to allow the results to be used for several purposes, thereby providing information that can be used to manage operational actions and optimize mining plans or engineering projects. In this work we analysed monitoring data (pore pressure and displacement) and its correlation with the tension and displacement of the mass of an established failure slope calculated using the finite element method. To optimize the back-analysis, a Python language routine was developed using input data (point coordinates, parameter matrix, and critical section) to use software with the rock mass parameters (cohesion, friction angle, Young’s modulus, and Poisson’s ratio). For the back-analysis, the Mohr-Coulomb criterion was applied with the shear strength reduction technique to obtain the strength reduction factor. The results were consistent with both the measured displacements and the maximum deformation contours, revealing the possible failure mechanism, allowing the strength parameters to be calibrated according to the slope failure conditions, and providing information about the contribution of each variable (parameter) to the slope failure in the study area.

Keywords open pit; slope failure prediction; inverse velocity technique; slope monitoring data; back-analysis.

Introduction The extraction of minerals by open pit mining is widely practised throughout the world. In some cases, large mines with very high and steep slopes are designed to allow economic gains. However, such designs can also promote high-risk situations, such as disastrous instabilities, with critical social, economic, and environmental consequences, especially in events culminating in the loss of human life.

Thus, a balance between operational security and mining economics must be sought. In this context, geotechnical studies play an important role.

During the mining process, excavation activities modify the initial stresses on the rock mass. In an effort to re-equilibrate, these stresses can cause instability, thereby enhancing the possibility of slope failures on the bench scale, inter-ramp scale, and/or mine scale. Geotechnicians must therefore consider (mainly during the operational phase) a variety of factors that can contribute to instabilities, such as increases in shear stresses with the removal of lateral support (resulting in erosion, falling blocks, and subsidence), changes in the groundwater level and corresponding increases in pore pressure, overloading of slopes, rainfall, external vibrations, and natural stress relief mechanisms with movements/displacements of the slope.

Rock mass behaviour during mining can be assessed through geotechnical monitoring with radars, topographic prisms, piezometers, and water level indicators, all of which allow the acquisition of deformation/displacement and pore pressure data and enable the water table position to be identified. To date, with technological advances and automation, these techniques have strengthened the reliability of the acquired data and increased the speed of analysis.

In the back-analysis of failure events, the monitoring data (displacements, pore pressures, and water table levels) can be correlated to calibrate the strength parameters. This approach ensures detailed analyses in geotechnical studies. Moreover, the mining geometry can be optimized, which may result in operational and financial gains for the company.

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Correlations of geotechnical monitoring data in open pit slope back-analysis Geotechnical monitoring Cawood and Stacey (2006) stated that slope monitoring can provide early warnings of disasters to avoid loss of life, damage to equipment, reduced production, and possibly mine closure.

These considerations were subsequently corroborated by Severin et al. (2014).

According to Carlà et al. (2017), the detection of slope deformation, which can cause slope failure, is a critical aspect in the fields of geomechanics and engineering geology. However, mitigating the risk of slope failure requires knowledge of various aspects of structural geology, including the properties of the rock mass and the influences of water and other external forces in the monitored area. Severin et al. (2014) reported that slope monitoring using radars, geodetic prisms, visual observations, and other methods constitutes a key component in modern risk management programmes for most mining companies.

Moore, Imrie, and Baker (1991) noted that effective slope monitoring should incorporate two stages: investigation monitoring and forecast monitoring. Investigation monitoring should provide an understanding of the typical behaviour and response of a slope over time to external events (rainfall and seasonal variations), while forecast monitoring should generate warnings of changes in behaviour, allowing damage to be delimited or enabling interventions to be implemented to prevent dangerous landslides.

According to Eberhardt, Watson, and Loew (2008), imminent failures are prevented by using phenomenological (temporal prediction) techniques that are based on measurements of surface displacements that, if analysed and extrapolated, yield accelerations that may exceed a previously established limit based on published standards.

In addition, Fukuzono (1985) proposed an inverse method that consists of calculating the rate of deformation (velocity) of a slope and plotting the inverse of the rate of deformation as a time function (inverse velocity versus time). This technique has been mentioned frequently in the literature (Rose and Hungr, 2007; Osasan and Stacey, 2014; Intrieri and Gigli, 2016). As the velocity or rate of deformation increases, the inverse will tend towards zero. With the use of linear regression, the intersection with the horizontal axis can be obtained; this intercept represents the expected time of collapse (Figure 1).

According to Venter, Kuzmanovic, and Wessels (2013), predictions of the expected collapse time are an important aspect of managing the stability of an open pit since such determinations enable appropriate actions to be taken. However, Severin et al. (2014) noted that false alarms resulting from data

Figure 1—Inverse velocity versus time prior to slope failure (Fukuzono, 1985)

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acquired from instrumentation and the uncertainties in this data constitute a problem. If an alarm is triggered and a failure occurs, the procedure is a success; otherwise, a false alarm can result in costs associated with stopping and delaying production in addition to a loss of faith in the system.

Stacey, Franca, and Beale (2018) postulated that the main differences among the applications of various monitoring methods can be attributed to the types of failures that can be anticipated. For example, hard, fractured rocks are predisposed to the rapid development of failures and rockfalls, whereas weak rocks fail more slowly, reflecting the occurrence or initiation of plastic deformation.

Most mining companies have modernized their conventional monitoring systems though automation and/or implementing advanced technologies such as radars (terrestrial radars and interferometric synthetic aperture radar (InSAR), global positioning systems (GPS, global navigation satellite systems (GNSS)), robotic stations, piezometers (electric, vibrating wire, pneumatic), and other instruments that reduce the time needed for data acquisition and boast greater reliability. In addition, these updated systems do not expose people to areas of risk (Benoit et al., 2015; Brown, Kaloustian and Roeckle, 2007; Atzeni et al., 2015; Dick et al., 2015; Stacey, Franca, and Beale, 2018; Carlà et al., 2018).

Case study The investigated slope is situated on the east wall of the open pit.

In 2017, a failure occurred, involving approximately four benches and the mass movement of material to lower benches. The affected area is shown within the yellow rectangle in Figure 2.

Geological aspects The study area is located within the geological domain of the Parauapebas and Carajás formations. The former is composed of generally altered metabasic volcanic rocks (basalts and diabases), while the latter contains ore represented by intercalated thick layers of jaspilites and lenses of soft and hard haematite (iron ore) cut by dykes and sills of basic rock. The main geological and structural properties of the eastern sector of the open pit are dominated by a shear zone of ductile to ductile-brittle character with a general strike ranging from E-W to NE-SW. This shear zone constitutes a weak surface that forms the contact

Figure 2—Distribution of instruments and location of section A-A The Journal of the Southern African Institute of Mining and Metallurgy

Correlations of geotechnical monitoring data in open pit slope back-analysis between the haematite formation and the metabasic rocks.

Dykes (discordant, with a thickness reaching 20 m and striking approximately N and NE) and a sill (with a thickness ranging from 0.5 m to 3.0 m) also occur. The sill underwent a significant failure event in 2017.

Hydrogeological aspects The groundwater level was efficiently depleted by the mine drainage system. In metabasic and altered rocks (mafic (MS), semi-decomposed mafic (MDS), and the decomposed mafic sill (MD)), the rock mass was assumed to be partially drained, a situation that occurs at a certain distance from the slope surface where the groundwater level drops in advance of the excavation, mainly by the escape of fluids through fractures.

identify possible measurement errors and/or noise. In the case of the rate of displacement (velocity), filtering was performed through linear regression analysis, as suggested by Rose and Hungr (2007).

The results of the displacement monitoring are collected as time series, t1... ti, d1... di, where ti and di are the most recent time and displacement, respectively, and N is the number of observations. Thus, the current rate of displacement is given by: [1] where [2]

2017 geotechnical event During geotechnical inspections in February 2017, cracks were observed on the upper slope in the eastern area of the open pit, below the main ramp. The area is characterized by geological, structural, and geotechnical properties that increase the risk of failure events, especially during the rainy season. Thus, taking into account these aspects in conjunction with regular geotechnical inspections, which include a network of prisms, piezometers, and groundwater level sensors to identify risk indicators, a terrestrial radar was installed to gain a deeper understanding of the deformational behaviour of the slope. This area presents geological structural and geotechnical conditions that can lead to instability, with some events having occurred in previous years (2007 and 2013).

The installed instrumentation showed variations in the measurements, mainly in the radar. However, the deformation/ displacement process accelerated only 24 hours before the collapse, which occurred at 08:50 on 12 March 2017.

The failure mode was considered planar/circular, with the mechanism developed in the contact between friable materials, the friable haematite (HF) with high permeability, and MD that exists along the entire eastern flank, which is a very low permeability material. The failure surface extended over a height of approximately 60 m. Figures 3 and 4 show the situation before failure and after failure, respectively.

[3] For the radar data, no filtering of the variables (deformation, velocity and inverse velocity) was performed as this data was extracted directly from the results produced by the software.

Thus, the probable periods of collapse were established through analyses of graphs, and the findings were subsequently compared with actual collapse data. Table I describes the instruments used in the back-analysis together with the measurement period.

Figure 2 illustrates the location of the instruments on the east flank of the central pit and the section utilized for the backanalysis. The differences in the topography before (28/02/2017) and after (30/03/2017) the event were considered to define the failure surface with the highest possible precision.

In this work, rainfall was considered an equal contributor with the structural conditions in the area (MD). Although the pore pressure varied during the rainy season, the water table was fixed in the back-analysis, and the last measurement obtained with the closest piezometer to the affected region was the level selected to calibrate the strength and elasticity parameters. The recorded pore pressure value was 45.202 mH2O at 08:27 on 13/03/2017, with a groundwater level of 646.901 m.

Methodology Data collection and analysis The information applied in this study is based on failure histories, topographic surveys, geotechnical mapping, geological (structural) data, laboratory testing, geotechnical monitoring (piezometric, groundwater level, prism, and surface radar data), and rainfall data. For the prism data, the horizontal displacement vector was utilized in the analysis. The data was filtered to

Figure 4—Open pit after failure (March 2017)

Table I

Analysed instrumentation and monitoring period Instrumentation

Figure 3—Open pit before failure (February 2017) The Journal of the Southern African Institute of Mining and Metallurgy

Prisms Piezometers Radar

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Analysed period

Parameter

12/01/2017–13/03/2017 01/12/2017–30/04/2017 20/02/2017–30/04/2017

Displacement Pore pressure Deformation

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Correlations of geotechnical monitoring data in open pit slope back-analysis The Mohr-Coulomb failure criterion was employed in the back-analysis, and the materials within the failure surface (HF and MD) were considered perfectly plastic; the other materials were considered elastic with no variations in their parameters.

The back-analysis was performed using RS2 9.0 (Rocscience) software, and the parameters were determined through the finite element method (FEM) to calculate the stresses and displacements in the rock mass using the shear strength reduction (SSR) technique (Matsui and San, 1992) to identify the strength reduction factor (SRF). The critical SRF is defined as the point when the analysis fails to converge and the deformation increases very quickly. Accordingly, the strength parameters (Table II) were obtained from laboratory tests on samples collected in the mine in campaigns from years prior to this work.

The calibration limits (Table III) were established in accordance with the maximum and minimum values obtained from the database of laboratory test results in campaigns prior to this study.

For this study, a routine was written in the Python programming language to assess n sets of variables, namely, the c (cohesion), E (friction angle), E (Young's modulus), and v (Poisson's ratio) of MD and HF. These variables, (which were calibrated n times), were chosen based on their important role in the failure process and to provide the displacement responses of the radar pixels within the line of sight of the equipment. The Python routine uses the X, Y, and Z coordinates of the top and bottom positions of the section in addition to the coordinates of the pixels and position of the radar.

A matrix was generated as the input with 100 combinations of the cohesion, friction angle, Young's modulus, and Poisson’s ratio. The predicted displacements were compared with the measured displacements. This process was repeated until, for a given set of input parameters, the predicted displacements approached the measured displacements with an acceptable error of approximately 5%.

To calibrate the displacements, the pixels imaged along the critical section (Figure 2) were chosen, and their behaviour (trends, cycles, and oscillations) analysed according to the date available prior to the failure. The radar operates using the phase variation between two scans. These measurements are applied to calculate the displacement of the slope surface; the displacement vector is not used for prism monitoring. Table IV presents information about the pixel size, distance to the radar, and values measured at a time near collapse.

Figure 5 presents the critical section modelled with the materials in the study area, namely the data obtained through geotechnical mapping and field observations. The failure surface, with a thickness of approximately 2 m, is defined, involving the

contact between the HF and the MD, with the latter playing a significant role in the failure process. The contact between the materials was modelled taking into account the geological model, geotechnical mapping, and field observations after failure.

Results and discussion Geotechnical monitoring data During the rainy season of 2017 (November 2016 to April 2017), the total rainfall reached 1311.10 mm, with a maximum of 920.5 mm before failure. The high rainfall resulted in recharge and saturation in the MD. The rain that fell on the previous days (a daily amount of 72 mm) contributed to the triggering of the failure, enabling movement accelerations on 11/03/2017 and 12/03/2017, as shown in the graph in Figure 6.

Event prediction was performed using the inverse velocity technique proposed by Fukuzono (1985). This method establishes that the time needed to accelerate the failure, while taking into account the gravitational load, is inversely proportional to the rate of deformation (velocity). In this analysis, the date corresponds to a period of 24 hours, which initially did not show a tendency towards imminent failure until 03/11/17.

The deformation was continuous, however, without progressive acceleration.

Figure 7 demonstrates that the deformation started to change at approximately 05:00 on 03/12/17, with a tendency to collapse between 08:49 and 11:30, a forecast very close to the actual time of failure.

Table III

Calibration limits Material MD HF

c (MPa)

f (°)

E (MPa)

v

[0.098; 0.085] [0.123; 0.114]

[26; 18] [42; 38]

[244; 240] [450; 445]

[0.23; 0.29] [0.177; 0.21]

Table IV

Pixel details and maximum measured values Pixel A3 A4 A6 A8

Size (m)

Radar distance (m) Deformation (mm)

10.28 ×× 10.24 10.14 × 10.11 9.97 × 9.93 9.96 × 9.92

1180.21 m 1164.53 m 1144.75 m 1143.39 m

97.869 127.864 156.281 140.681

Date 12/03/2017 12/03/2017 12/03/2017 12/03/2017

Table II

Strength parameters utilized in the back-analyses Lithology Mafic Semi-decomposed mafic Decomposed mafic sill Friable haematite Compact jaspilite Canga Shear zone

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g (dry) (MN/m3)

g (sat) (MN/m3)

Cohesion (MPa)

f (°)

E (MPa)

v

29 30 18.5 37 37 30 19

29 30 20 38 37 30 20

3.20 0.24 0.09 0.12 3.75 0.07 1.00

50 36 26 38 48 43 18

70 080.0 5 719.3 244.9 450.6 92 080.0 22 280.0 -

0.250 0.300 0.272 0.289 0.177 0.245 -

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Correlations of geotechnical monitoring data in open pit slope back-analysis

Figure 5—Localized pixels along the critical section

Figure 6—Displacement stages

Figure 7—Deformation measurements, inverse velocity technique, and collapse prediction The Journal of the Southern African Institute of Mining and Metallurgy

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Correlations of geotechnical monitoring data in open pit slope back-analysis The deformation rates in this short period were so high that, in less than 4 hours from the onset of acceleration, the rock mass lost its strength, causing the slope to fail.

The monitoring data showed precise values. Consequently, analysis of the data enabled not only an intervention in the study area but also a back-analysis to evaluate the correlation and the deformational behaviour of the slope. Table V shows the resulting maximum and average values for all the monitoring data from the failure.

Back-analysis

The results of a series of back-analysis calculations showed SRF values varying between 1.11 and 0.87. The maximum displacements were concentrated near SRF = 0.95.

The cohesion, friction angle, Young's modulus, and Poisson’s ratio for materials HF and MD varied; these parameters remained constant for the other materials present in the modelled section.

The analysis showed that seven series had displacement values similar to those measured in the area (pixels) at the moment of collapse (Tables VI and VII). According to the acceptability criterion defined for this study, which establishes ±5% variation between the measured displacements and the calculated displacements, the series with IDs 78, 81, 88, and 89 were

discarded. Therefore, the values calculated for pixel A3 were not acceptable.

For the remaining three series, the model was calibrated with the set of parameters containing the smallest variation between the measured displacements and the calculated displacements using the pixels at the base of the slope (in this region, the stress concentrations are higher, and in this case, allow greater displacements). Thus, the set of parameters for ID = 87 presented these conditions (Table VIII).

The results obtained from the back-analysis and statistical data show that the variables (parameters) with the greatest contributions to the failure were the cohesion, friction angle, and Poisson's ratio, which influenced the MD (approximately 93%) more than the HF (Figure 8).

The maximum deformations obtained at the time of failure were calculated from the calibrated parameters, considering that the failure mechanism developed along the contact between the MD and the HF. The maximum displacements were observed at pixel A6, with a calculated displacement of 153.79 mm and measured displacement of 156.281 mm, thereby defining the region with the greatest acceleration. The collapse originated in this region.

Table V

Statistical results of geotechnical monitoring Instrument

Parameter

Unit

V alues from November 2016 to the day of collapse

Average Rainfall station

Maximum

Maximum value date

Rainfall

mm

9.3

920.5

12/03/2017

Elevation Pore pressure

m mH20

642.158 40.458

646.918 45.218

12/03/2017 08:27 12/03/2017 08:27

XY Z Vel. XY Vel. Z

mm mm mm/day mm/day

21.317 9.530 0.486 0.263

117.200 35.100 2.934 0.755

09/03/2017 11:34 09/03/2017 11:34 09/03/2017 11:34 09/03/2017 11:34

Pixel A3

Deformation (PS) Velocity (PS) Deformation (SS) Velocity (SS) Deformation (TS/C) Velocity (TS/C)

mm mm/day mm mm/day mm mm/day

4.148 0.740 8.580 2.667 37.185 15.145

8.392 7.523 21.624 14.197 97.869 64.057

12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50

Pixel A4

Deformation (PS) Velocity (PS) Deformation (SS) Velocity (SS) Deformation (TS/C) Velocity (TS/C)

mm mm/day mm mm/day mm mm/day

4.237 0.541 6.003 1.193 32.972 22.581

10.266 15.604 15.491 17.831 127.864 104.349

12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50

Pixel A6

Deformation (PS) Velocity (PS) Deformation (SS) Velocity (SS) Deformation (TS/C) Velocity (TS/C)

mm mm/day mm mm/day mm mm/day

4.021 0.419 4.343 0.279 40.226 16.258

9.292 8.498 11.715 10.082 156.281 131.650

12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50

Pixel A8

Deformation (PS) Velocity (PS) Deformation (SS) Velocity (SS) Deformation (TS/C) Velocity (TS/C)

mm mm/day mm mm/day mm mm/day

4.019 0.500 12.493 2.952 51.994 25.446

9.210 8.620 25.242 15.252 140.681 98.692

12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50 12/03/2017 08:50

PZCV-02

Prism: P28

PS: Primary stage, SS: Secondary stage, TS/C: Tertiary stage/collapse

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Correlations of geotechnical monitoring data in open pit slope back-analysis Table VI

Strength reduction factor ID

78 81 82 83 87 88 89

c (MPa)

f (°)

0.114 0.114 0.114 0.114 0.114 0.114 0.114

38 38 38 38 38 38 38

HF

MD (Sill) SRF

E (MPa)

v

c (MPa)

f (°)

E (MPa)

v

446.0 445.5 446.5 447.0 446.5 446.5 447.0

0.18 0.18 0.19 0.19 0.19 0.18 0.18

0.085 0.085 0.085 0.085 0.088 0.085 0.085

22 22 22 22 21 22 22

243.0 242.5 243.5 243.0 243.5 243.5 243.0

0.28 0.28 0.30 0.30 0.30 0.28 0.28

0.94 0.94 0.94 0.94 0.95 0.94 0.94

Table VII

Back-analysis results ID

A3 (97.869 mm)

A4 (127.864 mm)

A6 (156.281 mm)

A8 (140.681 mm)

SRF

78 81 82 83 87 88 89

106.69 106.75 101.34 98.31 101.36 107.26 105.42

131.21 131.76 130.19 126.34 130.27 132.12 130.27

153.61 153.60 153.24 149.27 153.79 154.78 152.65

140.31 140.40 139.83 135.82 139.84 141.38 139.27

0.94 0.94 0.94 0.94 0.95 0.94 0.94

( ) monitored values

Figure 8—Contributions of the parameters in the instability

Table VIII

Back-analysis parameters

c (MPa)

0.114

f (°) 38

HF E (MPa)

v

c (MPa)

446.5

0.19

0.088

MD (Sill) f (°)

E (MPa)

v

21

243.5

0.30

HF: Friable haematite MD: Decomposed mafic sill

Figure 9 presents the critical section analysed with the calibrated parameters, failure surface defined in the sill, and area where the displacements initiated the collapse.

Conclusions Rainfall played an important role in the failure event and led to mass saturation and lubrication of the HF/MD sill contact and The Journal of the Southern African Institute of Mining and Metallurgy

pre-existing cracks. These processes were pronounced due to the ineffectiveness of the operational drainage system. The rainfall and failure mechanism already established in the flank of the pit acted simultaneously to trigger the event.

The back-analysis using the SSR technique revealed a good correlation between the calculated data and the data measured with ground radar, and the evaluated parameters were acceptably estimated. In this study, the prism data alone was not sufficient for the analysis, but the use of the SSR technique with this data (vectors) could provide more effective parameter calibration than that with ground radar data. Regardless of the type of monitoring, the SSR technique is relevant for calibrating the strength parameters and providing information about the development of the failure mechanism. This information will be particularly helpful for geotechnicians in the evaluation of future projects in the area.

The statistical analysis of the contributions of the strength and elasticity parameters from the back-analysis was not VOLUME 121

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563 563 ◀

Correlations of geotechnical monitoring data in open pit slope back-analysis

Figure 9—Maximum deformations in the calibrated section

conclusive in relation to Poisson’s ratio and Young’s modulus for MD sill material. Notably, the divergent representativeness of the results suggests that more detailed studies of these parameters for this type of material are needed for similar failure events.

In general, a circumspect analysis of geotechnical monitoring data serves as a tool for better understanding the deformational behaviour of a slope and provides a more assertive and preventive approach to avoid failures, or if a failure is imminent, to reduce the impacts.

Acknowledgements We thank Vale S.A. for making the data available and Institute Technologic Vale for guiding and supporting this work.

Dick, G.J., Eberhardt, E., Cabrejo-Liévano, A.G., Stead, D., and Rose, N.D. 2015.

Development of an early-warning time-of-failure analysis methodology for open-pit mine slopes utilizing ground-based slope stability radar monitoring data. Canadian Geotechnical Journal, vol. 52, no 4. pp. 515–529.

Eberhardt, E., Watson, A.D., and Loew, S. 2008. Improving the interpretation of slope monitoring and early warning data through better understanding of complex deep seated landslide failure mechanisms. Landslide and Engineering Slopes: From the Past to the Future. Chen, Z., Zang, J., Li, Z, Wu, F., and Ho, K. (eds).

Taylor and Francis, London.

Fukuzono, T. 1985. A new method for predicting the failure time of a slope.

Proceedings of the Fourth International Conference and Field Workshop on Landslides. Japan Landslide Society, Tokyo. pp. 145–150.

Contributions

Intrieri, E. and Gigli, G. 2016. Landslide forecasting and factors influencing predictability. Natural Hazards and Earth System Sciences, vol. 16, no 12. doi:10.5194/nhess-16-2501-2016

Aristotelina Ferreira Silva carried out the study.

Juan Manuel Girao Sotomayor wrote the manuscript and supervised the study.

Vidal Félix Navarro Torres supervised the study.

Matsui, T., and San, K C. 1992. Availability of shear strength reduction technique.

Proceedings of the Conference on Stability and Performance of Slopes and Embankments II, Berkley, CA. American Society of Civil Engineers, New York. pp. 445–460.

References Atzeni C., Barla M., Pieraccini M., and Antolini F. 2015. Early warning monitoring of natural and engineered slopes with ground-based synthetic-aperture radar.

Rock Mechanics and Rock Engineering, vol. 48, no.1. pp. 235-246.

Benoit, L., Briole, P., Martin, O., Thom, C., Malet, J.P., and Ulrich, P. 2015. Monitoring landslide displacements with the Geocube wireless network of low-cost GPS. Engineering Geology, vol. 195, pp. 111–121.

Brown, N., Kaloustian, S., and Roeckle, M. 2007. Monitoring of open pit mines using combined GNSS satellite receivers and robotic total stations. Proceedings of the International Symposium on Rock Slope Stability in Open Pit Mining and Civil Engineering. Perth. Potvin, Y. (ed.). Australian Centre for Geomechanics, University of Western Australia, Nedland. pp. 417–429.

Carlà, T., Farina, P., Intrieri, E., Botsialas, K., and Casagli, N. 2017. On the monitoring and early warning of brittle slope failures in hard rock masses: Examples from an open-pit mine. Engineering Geology, vol. 228. pp. 71–81.

Carlà, T., Farina, P., Intrieri, E., Ketizmen, H., and Casagli, N. 2018. Integration of ground-based radar and satellite InSAR data for the analysis of an unexpected slope failure in an open-pit mine. Engineering Geology, vol. 235. pp. 39–52.

Cawood, F.T. and Stacey, T.R. 2006. Survey and geotechnical slope monitoring considerations. Journal of the South African Institute of Mining and Metallurgy, vol. 106, no. 7. pp. 495–501.

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Moore, D.P., Imrie, A.S., and Baker, D.G. 1991. Rockslide risk reduction using monitoring. Proceedings of the Canadian Dam Association Meeting, Whistler, BC. National Research Council Canada. pp. 245–258.

Osasan, K.S. and Stacey, T.R. 2014. Automatic prediction of time to failure of open pit mine slopes based on radar monitoring and inverse velocity method.

International Journal of Mining Science and Technology, vol. 24, no 2. pp. 275–280.

Rose, N.D. and Hungr, O. 2007. Forecasting potential rock slope failure in open pit mines using the inverse-velocity method. International Journal of Rock Mechanics and Mining Sciences, vol. 44. pp. 308–320.

Severin, J., Eberhardt, E., Leoni, L., and Fortin, S. 2014. Development and application of a pseudo-3D pit slope displacement map derived from ground-based radar.

Engineering Geology. vol. 181. pp. 202–211.

Stacey, P., Franca, P., and Beale, G. 2018. Design implementation and operational consideration. Guidelines for Open Pit Slope Design in Weak Rocks. Vol. 1.: Martin, D.and Stacey, P. (eds). CSIRO Publishing, Clayton, Australia.

Venter, J., Kuzmanovic. A., and Wessels, S. 2013. An evaluation of CUSUM and inverse velocity methods of failure prediction based on two open pit instabilities in the Pilbara. Slope Stability 2013: Proceedings of the 2013 International Symposium on Slope Stability in Open Pit Mining and Civil Engineering, Brisbane, Australia. Dight, P.M. (ed.). Australian Centre for Geomechanics, Perth. pp. 1061-1076. https://doi.org/10.36487/ACG_rep/1308_74_Venter u The Journal of the Southern African Institute of Mining and Metallurgy

NATIONAL & INTERNATIONAL ACTIVITIES 2021 11 November 2021 — 17TH Annual Online Student Colloquium 2021 Contact: Gugu Charlie Tel: +27 11 834-1273/7 Fax: +27 11 838-5923/833-8156 E-mail: gugu@saimm.co.za Website: http://www.saimm.co.za 15, 17, 19, 22, 24, 26 November 2021 — Global Tailings Standards and Opportunities Online Conference 2021 ‘For the Mine of the Future’ Contact: Gugu Charlie Tel: +27 11 834-1273/7 Fax: +27 11 838-5923/833-8156 E-mail: gugu@saimm.co.za Website: http://www.saimm.co.za

2022 23–24 February 2022 — Drill and Blast Hybrid Short Course 2022 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 February 27–March 3, 2022 — TMS Furnace Tapping 2022 Anaheim Convention Center & Anaheim Marriott , Anaheim, California, USA https://www.tms.org/AnnualMeeting/TMS2022/Programming/Furnace_Tapping_2022/AnnualMeeting/TMS2022/Programming/furnaceTapping.aspx?hkey=718f6af7-1852-445cbe82-596102913416 20–27 May 2022 — ALTA 2022 Hybrid Conference 2022 Perth, Australia, Tel: +61 8 9389 1488 E-mail: alta@encanta.com.au Website: www.encanta.com.au 14–16 June 2022 — Water | Managing for the Future Vancouver, BC, Canada https://www.mineconferences.com 9–13 July 2022 — Sustainable Development in the Minerals Industry (SDIMI) 2022 10th International Hybrid Conference Swakopmund Hotel and Entertainment Centre, Swakopmund, Namibia Contact: Gugu Charlie Tel: +27 11 834-1273/7 Fax: +27 11 838-5923/833-8156 E-mail: gugu@saimm.co.za Website: http://www.saimm.co.za

The Journal of the Southern African Institute of Mining and Metallurgy

21–25 August 2022 — XXXI International Mineral Processing Congress 2022 Melbourne, Australia + Online www.impc2022.com 24–25 August 2022 — Battery Materials Conference 2022 Misty Hills Conference Venue, Muldersdrift, Johannesburg, South Africa 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 8–14 September 2022 — 32nd SOMP Annual Meeting and Conference 2022 Windhoek Country Club & Resort, Windhoed, Namibia 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 28–30 September 2022 — PGM The 8th International Conference 2022 Sun City, Rustenburg, South Africa 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 28–29 September 2022 — Thermodynamic from Nanoscale to Operational Scale (THANOS) International Hybrid Conference 2022 on Enhanced use of Thermodynamic Data in Pyrometallurgy Teaching and Research Mintek, Randburg, South Africa 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 24–27 October 2022 — 8th Sulphur and Sulphuric Acid Conference 2022 The Vineyard Hotel, Newlands, Cape Town, South Africa 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

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Company affiliates The following organizations have been admitted to the Institute as Company Affiliates Company Affiliate

Expectra 2004 (Pty) Ltd

Ltd

3M South Africa (Pty) Limited

Exxaro Coal (Pty) Ltd

MSA Group (Pty) Ltd

AECOM SA (Pty) Ltd

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Multotec (Pty) Ltd

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Filtaquip (Pty) Ltd

Murray and Roberts Cementation

African Pegmatite (Pty) Ltd

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G H H Mining Machines (Pty) Ltd

Northam Platinum Ltd - Zondereinde

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Opermin Operational Excellence

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Optron (Pty) Ltd

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The Journal of the Southern African Institute of Mining and Metallurgy

8th

SULPHUR AND SULPHURIC ACID CONFERENCE | 2022 24 OCTOBER 2022 - WORKSHOP

Sulfuric Acid Catalysis - Key Parameters to Increase Efficiency and Lower Costs

25-26 OCTOBER 2022 - CONFERENCE 27 OCTOBER 2022 - TECHNICAL VISIT THE VINEYARD HOTEL, NEWLANDS, CAPE TOWN, SOUTH AFRICA

BACKGROUND The production of SO2 and sulphuric acid remains a pertinent topic in the Southern African mining and metallurgical industry, especially in view of the strong demand for, and increasing prices of, vital base metals such as cobalt and copper.

The electric car revolution is well underway and demand for cobalt is rocketing.

New sulphuric acid plants are being built, comprising both smelters and sulphur burners, as the demand for metals increases. However, these projects take time to plan and construct, and in the interim sulphuric acid is being sourced from far afield, sometimes more than 2000 km away from the place that it is required.

The need for sulphuric acid ‘sinks’ such as phosphate fertilizer plants is also becoming apparent.

All of the above factors create both opportunities and issues and supply chain challenges.

To ensure that you stay abreast of developments in the industry, the Southern African Institute of Mining and Metallurgy invites you to participate in a conference on the production, utilization, safe transportation and conversion of sulphur, sulphuric acid, and SO2 abatement in metallurgical and other processes, to be held in October 2022 in Cape Town.

FORMAT OF THE EVENT At this point in time, the event is planned as a full contact conference with international participation through web links. It is also planned to hold technical visits to nearby facilities.

The situation will be constantly reviewed, and if it appears that the effects of the pandemic are still such as to pose a threat to the health and safety of delegates, this will be changed to a digital event.

OBJECTIVES •

•

•

•

To expose delegates to issues relating to the generation and handling of sulphur, sulphuric acid, and SO2 abatement in the metallurgical and other industries.

Provide an opportunity to producers and consumers of sulphur and sulphuric acid and related products to be introduced to new technologies and equipment in the field.

Enable participants to share information about and experience in the application of such technologies.

Provide an opportunity for role players in the industry to discuss common problems and their solutions.

WHO SHOULD ATTEND The Conference will be of value to: Metallurgical and chemical engineers working in the minerals and metals processing and chemical industries Metallurgical/chemical/plant management Project managers Research and development personnel Academics and students Technology providers and engineering firms Equipment and system providers Relevant legislators

EXHIBITION AND SPONSORSHIP There are a number of sponsorship opportunities available. Companies wishing to sponsor or exhibit should contact the Conference Co-ordinator.

FOR FURTHER INFORMATION, CONTACT: Camielah Jardine, Head of Conferencing E-mail: camielah@saimm.co.za Web: www.saimm.co.za

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