Saimm 202507 jul

VOLUME 125 NO. 7 JULY 2025

The Southern African Institute of Mining and Metallurgy

OFFICE BEARERS AND COUNCIL FOR THE 2024/2025 SESSION

President E. Matinde

President Elect

G.R. Lane

Senior Vice President

T.M. Mmola

Junior Vice President

M.H. Solomon

Incoming Junior Vice President

S.J. Ntsoelengoe

Immediate Past President

W.C. Joughin

Honorary Treasurer

W.C. Joughin

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W. Broodryk M.C. Munroe

Z. Fakhraei S.M. Naik

B. Genc G. Njowa

K.M. Letsoalo S.M. Rupprecht

S.B. Madolo A.T. van Zyl

M.A. Mello E.J. Walls

K. Mosebi

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A.D. Coetzee

L.T. Masutha

Past Presidents Serving on Council

N.A. Barcza C. Musingwini

R.D. Beck S. Ndlovu

Z. Botha J.L. Porter

V.G. Duke M.H. Rogers

I.J. Geldenhuys G.L. Smith

R.T. Jones

G.R. Lane – TP Mining Chairperson

Z. Botha – TP Metallurgy Chairperson

K.W. Banda – YPC Chairperson

C.T. Chijara – YPC Vice Chairperson

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S.O. Bada

P. den Hoed

I.M. Dikgwatlhe

M. Erwee

B. Genc

A.J. Kinghorn

D.E.P. Klenam

D.F. Malan

D. Morris

P.N. Neingo

S.S. Nyoni

M. Phasha

P. Pistorius

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N. Rampersad

Q.G. Reynolds

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VOLUME 125 NO. 7JULY 2025

Contents

Journal Comment: Finding the needle in the haystack by Q.G. Reynolds iv

President’s Corner: South Africa’s green hydrogen strategy: Challenges and opportunities by E. Matinde ..................................................................

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ISSN 2225-6253 (print) . ISSN 2411-9717 (online)

PROFESSIONAL TECHNICAL AND SCIENTIFIC PAPERS

Fatigue performance improvement using laser shock peening in high strength ductile metallic materials by S.L. George, R. Tait, A. Becker, A. Carlisle

Surface treatments like laser shock peening (LSP) significantly increase the potential impact on materials prevalent in the automotive and aerospace sectors. Through extensive fatigue testing and fractography investigations, it was determined that LSP resulted in a fatigue life increase of over five times for aluminium alloy AA7075-T6, 1.7x for titanium Ti-6Al-4V, and was also noted in AISI316 stainless steel. While fatigue life in all cases improved, the efficacy of LSP was shown to be material dependent.

Comparison of remote sensing techniques in mapping hydrothermal alteration associated with Ovacik epithermal gold-silver mineralization by C. Kwang, H. Uygycgil

There is a need to explore the potential of new robust remote sensing and geographic information system techniques in the field of mineral exploration. In this study, the argillic alteration associated with epithermal gold-silver mineralisation was mapped. Of the two methods used in mapping out the argillic alteration zones associated with gold-silver mineralisation, the Crósta technique produced the optimal argillic alteration. The study demonstrates the analytical capabilities of Google Earth Engine in image processing and analysis for mineral exploration.

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch as binder by

A.M. Garbers-Craig, I. Cameron, S. Ramjee

The cause of reduced workability and increased ageing of a blast furnace taphole clay was examined. The workability and Marshall extrusion pressure of the clay samples were evaluated to assess the extent of ageing of the clay. The analyses confirmed a chemical interaction between the resole resin and liquid pitch. Upon ageing of the resin-pitch mixtures, the resin underwent premature cross-linking, causing earlier onset of curing. This was identified as the primary cause of reduced workability, accelerated ageing, and increased Marshall extrusion pressure of the taphole clay.

The use of soundless chemical demolition agents in large scale in situ rock breaking applications in the mining industry by I. Maubane, P.L. Ngwenyama

Soundless chemical demolition agents (SCDAs) are proving to be the future of sustainable and environmentally-friendly mining. This study evaluated whether soundless chemical demolition can break large volumes of in situ rock. This was done by conducting five trials using Nex-Pand soundless chemical demolition agents in four different surface mining sites. These trials proved that soundless chemical demolition agents can replace explosives in areas closer to sensitive infrastructure and communities. This will ultimately enable mines to unlock value and access deposits in these areas.

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation by O.D. Eniowo, M. Onifade, H. Grobler, A.F. Mulaba-Bafubiandi, O.S. Otuogbai

This study looks at how artisanal and small-scale miners can attract formal funding, which will help them take the leap from artisanal state to a sustainable small-scale mining operation. This study evaluates the creditworthiness of artisanal and small-scale mining operators in Nigeria, using key credit risk parameters identified in the literature and through primary investigation. The results show that being a member of a mining cooperative society improves the chances of access to formal financing for miners within the sample group.

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data by T. Kgarume, M. van Schoor, M. Mpofu, H. Grobler

The South African mining industry has committed to achieving a state of zero harm for its workforce. Among the major safety concerns are falls of ground, a leading cause of injuries and fatalities. This study presents a comprehensive georeferencing methodology. The potential for integrating diverse datasets to construct insightful models of the underground mining environment is illustrated. This potential is likely to offer a holistic understanding of the mining environment, providing essential information to decision-makers.

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Comment

IFinding the needle in the haystack

n the course of our engineering work on mining and metallurgical plants we are often called upon to evaluate the merits of different choices in process flowsheets, operating parameters and philosophies, raw material selection, and many others. The phenomenological complexity of the minerals industry usually means that each of these aspects is parameterised by a large number of variables, and there are also strong coupling effects between them – one changes a feed-rate setting here, and even though it fixes the immediate production problem over here, it also affects several other things over there in ways that one did not expect.

In the digital age we have access to powerful process and systems models, which can create virtual analogues (or “digital twins” if one prefers catchy jargon) of our real-world plants, which can make decision-making easier. However, in the pursuit of improved accuracy these models often start to become as impenetrable and confusing a black box as the actual thing they are trying to simulate. This is especially true when data-centric artificial intelligence and machine learning methods are included in the mix. Manually exploring such systems models by making basic changes using oversimplified fundamental principles, can very quickly turn into an endless game of whack-a-mole to mitigate the cascade of unintended consequences.

To better manage this problem, two formal mathematical concepts are becoming increasingly useful as interface layers over complex systems models. Uncertainty quantification tracks the propagation of errors through a system from inputs to outputs, and sensitivity analysis identifies how strongly outputs are affected by changes in the inputs. In combination, these tools can help guide design or process optimisation studies to find small changes that yield large improvements while minimising undesirable side effects. They are well worth investigating to help us find the needles in our metallurgical haystacks.

Q.G. Reynolds

South Africa’s green hydrogen strategy: Challenges and opportunities President’s Corner

Significant global efforts have been dedicated to mitigate the man-made impacts of climate change and environmental degradation. However, recent policy shifts by some of the major economic jurisdictions, such as the United States, to focus on prioritising domestic economic growth, potentially at the expense of climate change, and stricter environmental regulation is a major cause for concern. For example, the US’s policy changes to truncate the roll-out of electric vehicle incentives and subsidies, among other roll-backs on clean energy transition initiatives, will provide interim fiscal relief but may inadvertently undermine climate change mitigation measures in the long term. In addition, the emerging geoeconomic order and tariff regimes will also have a disproportionate impact on both the demand-side and supply-side of climate-neutral technology cooperation and financing mechanisms. Consequently, this will delay the deployment of nascent interventions required to reduce emissions from energy- and greenhouse gas intensive industries. Obviously, the impact is disproportionately high in critical but hard-to-abate industries such as cement clinker production, iron and steel manufacturing, and public transportation, of which their global CO2 emissions are estimated to be roughly 8%, 7-9%, and up to 14%, respectively.

Due to the high costs of technology development and complexity of integration into existing systems, global cooperation in the development and financing mechanisms for sustainable climate-friendly technologies, such as green hydrogen, carbon capture, and storage technologies, has immense benefits to mankind. Green hydrogen, in particular, is considered to be one of the most promising energy carriers with immense environmental benefits. According to the Green Hydrogen Organisation (https://gh2.org/what-green-hydrogen), green hydrogen is produced via an electrochemical process to split water into hydrogen and oxygen using renewable sources of energy. Several technologies to produce green hydrogen at scale have been developed and/or are at different stages of development and commercialisation. However, each technology regime is characterised by its own inherent challenges and opportunities in terms of interoperability, efficiency, costs and availability. With substantial support from industry, policymakers and the public, the roll-out of green hydrogen technologies has been touted as a gamechanger, with massive global strategic efforts being deployed towards technology development and establishment of infrastructure and special economic zones, supported by the promulgation of targeted industrial policies as well as support with tax incentives and subsidies.

As part of the dual drive to attain both economic and energy sovereignty, South Africa launched an ambitious green hydrogen strategy to leverage the country’s abundant renewable energy resources, both for domestic use and for export. The Hydrogen Society Roadmap was launched in 2021 to support the implementation of the country’s green hydrogen economy, and its implementation is anchored on the attainment of four strategic outcomes, viz, (1) creating an export market for the country’s green hydrogen and allied products, (2) greening the power generation, (3) decarbonising the transportation and heavy industries and, (4) localising the green hydrogen supply chains. According to the Hydrogen Society Roadmap (https://gh2.org/countries/south-africa), the country aims to produce approximately 500,000 tonnes per annum of green hydrogen by 2030, achievable by 10 GW of electrolysis capacity in the Northern Cape special economic zones by 2030, and up to 15 GW by 2040. In addition, the Hydrogen Roadmap targets the deployment of 100 hydrogen-powered buses and trucks by 2025 and up to 500 buses and trucks by 2030, with the opportunity to create and sustain up to 30,000 jobs annually by 2040. According to a report by National Business Initiative (https://www.nbi.org.za/green-hydrogen-presents-theopportunity-as-the-fuel-for-the-future/), South Africa has the potential to produce green hydrogen for USD1.60 per kg by 2030, one of the lowest costs worldwide. These ambitious energy transition targets are applaudable, despite the challenges to achieve them being the stated timelines and current technoeconomic landscape. In

President’s Corner (continued)

addition, there is an urgent need to revisit the assumptions used to formulate the stated impact targets, if one is to take into account the unprecedented number of bankruptcies by green hydrogen technology startups and established global companies due to inhibitive development costs and complexity of the associated systems and technologies.

Indeed, green hydrogen is going to be a game-changer due to its potential to drastically reduce greenhouse gas emissions and drive innovations in sustainable technologies. However, it is evident that the suite of technologies to produce green hydrogen are emerging technologies, which are disproportionately prone to failure due to the high costs of electrolyser technologies, complexity of integration, and intermittency of renewable energy storage systems, among other challenges. Synergistic to bottlenecks from the complex technology systems is the valley of death faced by green hydrogen technologies due to misallocation of financing, and the “Lindy effects” arising from sunk costs and perceived performance of established fossil fuel-based technologies. Although the challenges in the roll-out of green hydrogen technologies is a global phenomenon, reliance on imported technologies, high cost, and availability of climate finance, further increases vulnerability for countries in the global south, South Africa included.

I had the privilege of listening to a keynote address by the Chairperson of SAIMM Limpopo Branch, Mr Steven Zulu, at the branch event held at the University of Limpopo earlier this month. In his opening address, Mr Zulu highlighted the need to develop sustainable technologies based on endogenous technology learning capabilities as a sovereign strategy to mitigate against perpetuation of external dependency, a phenomenon he referred to as “technology colonialism”. In particular, Mr Zulu highlighted some basic implementation strategies to attain technology sovereignty, such as building sustainable R&D and technology development skills. Furthermore, he emphasised the importance of developing home-grown technology alternatives, drive down technology costs, and unlock the ability to reverse engineer existing technologies to suit the domestic market requirements.

In conclusion, there is no doubt that green hydrogen can be a game-changer for South Africa and the region. However, the attainment thereof, together with the Green Hydrogen Roadmap impact targets, risk being wishful thinking unless collective efforts are channelled towards intensifying technology development initiatives to drive down costs and reduce dependency on imported technologies and components. Most importantly, openminded approaches are required to take advantage of emerging geoeconomic dynamics, so as to establish genuine collaborations with all leading green hydrogen technology developers globally. As the SAIMM, we commit to continue supporting the dissemination of technical knowledge required to sustain the localisation of technology know-how in this highly contested domain.

.

E. Matinde President, SAIMM

Affiliation:

1Centre for Materials Engineering, Department of Mechanical Engineering, University of Cape Town, South Africa

2Department of Mechanical Engineering, University of Cape Town, South Africa

Correspondence to:

S.L. George

Email: sarah.george@uct.ac.za

Dates:

Received: 13 May 2021

Revised: 20 Aug. 2022

Accepted: 25 Apr. 2025

Published: July 2025

How to cite:

George, S.L., Tait, R., Becker, A., Carlisle, A. 2025. Fatigue performance improvement using laser shock peening in high strength ductile metallic materials.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 125, no. 7, pp. 347–354

DOI ID:

https://doi.org/10.17159/2411-9717/1619/2025

ORCiD: S.L. George

http://orcid.org/0000-0002-6084-6361

Fatigue performance improvement using laser shock peening in high strength ductile metallic materials

by S.L. George1, R. Tait2, A. Becker2, A. Carlisle2

Abstract

Surface treatments like laser shock peening (LSP) induce residual stresses penetrating up to 1 mm deep, surpassing traditional methods, thus significantly increasing the potential impact on materials prevalent in the automotive and aerospace sectors. Through extensive fatigue testing and fractography investigations, it was determined that LSP resulted in a fatigue life increase of over five times for aluminium alloy AA7075-T6 and titanium Ti-6Al-4V. Increased fatigue life was also noted in AISI316 stainless steel, which showed increases in fatigue life of 1.7x. While fatigue life in all cases improved, the efficacy of LSP was shown to be material dependent.

Keywords

laser shock peening, compressive residual stress, fatigue performance, AA7075-T6, Ti-6Al-4V, AISI 316

Introduction

Fatigue failure is one of the most common failure mechanisms in industry where cyclic loading is experienced, contributing to an estimated 80% of all mechanical service failures (ASM International, 2008). Many such failures result from fatigue cracks initiating at stress concentration sites found on the surface of components (ASM International, 2008). The fatigue life of a component comprises fatigue crack initiation as well as crack propagation, with initiation commonly being related to the surface condition of the component, such as surface finish and the presence of surface or sub-surface defects. To prolong the fatigue life of components, surface treatments can be utilised. Surface treatments aim to achieve one or more of the following: reduction of local stress concentrations, removal or neutralisation of existing defects, or reduction of tensile residual stress through the introduction of compressive residual stress (CRS), (Cui, 2002). Such residual stresses are introduced into the surface layers of the component to increase the number of cycles associated with the crack initiation stage These local residual stresses (normally compressive) counteract applied stresses (often tensile in nature), thereby delaying the onset of crack initiation. Residual compressive stresses at crack tips and surrounding material can reduce fatigue crack propagation rates, resulting in an overall increase in the fatigue life of the component (Montross, et al., 2002). One of the most effective and widely used surface modification processes is shot peening (SP), which is a cost-effective and robust process. Studies have been conducted into how the shot peening process affects the residual stress (Peyre, 1996; Hammond, 1990). Magnitudes of residual stresses produced by a shot peening process are at least as great as tensile strength of the material being peened (Metal Improvement Company, 2005). However, the SP process itself has limitations (Montross, et al., 2002; Kulekci, Esme, 2014), such as deterioration of surface finish and limited penetration depth of the compressive residual stresses. LSP CRS layer is typically 4-5 times deeper that the depth achieved through traditional SP (Gupta et el., 2017). The laser shock peening process introduces a slightly rougher surface finish to the sample, when compared to SP (Montross, et al., 2002). Surface roughness plays a critical role in the crack initiation stage of the fatigue testing (Montross, et al., 2002). Therefore, it can be important to eliminate surface roughness as a contributing factor to facilitate a direct comparison between laser shock peened (LSP) samples and non-laser shock peened (non-LSP) samples. The induced roughness can be removed before the peened components can be put into service, but Montross, et al. (2002) raised a concern regarding the offset of shot peen induced benefits through the removal of some of the residual stress in the surface layer in the context of SP where the CRS layer is relatively thin. Liu et al. (2017) determined that the benefit of improved surface quality from post LSP polishing may be considered as overriding the disadvantage of a reduced CRS layer thickness if the penetration depths are great enough.

Fatigue performance improvement using laser shock peening

There are increasing demands for lower operational costs, higher safety measures, and better lifetime performance characteristics in industry. As such, significant pressure has been placed on the manufacturing systems and surface processing technologies to produce components that are ’near flawless’ and require as few processing steps as possible before completion. One of the surface treatment techniques that has been developed recently, in response to these demands, is laser shock peening (LSP). LSP utilises high-speed and high-energy neodymium lasers to focus short-duration coherent energy pulses onto the surface of the component. A high amplitude stress wave propagates into the material, thereby causing the surface layer to yield in compression, and plastically deform.

This localised plastic deformation produces both strain hardening and residual compressive stress at the surface of the laser peened component and extends below the surface of the material (Clauer, 1997), up to a depth of approximately 1 mm (Clauer, 2001) or more. LSP is reported to have a limited minimal effect on the surface roughness of the material (Zhang, 2010; Schubnell, 2023) compared to SP. The extent of roughness changes is directly related to the laser process parameters and the use of a protective coating layer during the LSP process (Zhang, 2010; Schubnell, 2023). The LSP process can be applied to the finished surface of a part or before the final finishing step. Regions inaccessible to SP, such as small fillets and notches, can be treated by being considered for treatment using LSP. The application of LSP to parts where line-of-sight access is available is very straightforward, making it an easy process to use in industry (Holmes, 2013). LSP has been shown to significantly improve the fatigue performances of engineered components (Montross, et al., 2002; Ren et al.).

Over and above the benefits to fatigue life, laser shock peening can result in improved performance in wear, corrosion, and stress corrosion cracking for a variety of materials, (Montross, et al., 2002; Abdullahi, 2014). The fatigue life of stainless steel has been investigated by various researchers (Peyre, 1996; Bikdeloo, et al., 2020) and has been shown to improve the fatigue life of components. Bikdeloo et al. (2020) showed that AISI 316L stainless steel exhibits approximately 125% increase in fatigue life with single LSP layers and 170% with a second layer. Other materials, such as carbon steels, aluminium alloys, titanium alloys, and nickel-based super-alloys have also shown improved fatigue life through the application of LSP, where scanning parameters were investigated in detail (Abdullahi, et al., 2014; Adu-Gyamfi, et al., 2018; Sano, et al., 2006; Correa, et al., 2015]). Other considerations regarding the efficacy of LSP are related to crystal structure and stacking fault energies of the material, as deformation mechanisms differ, thus, affecting the CRS and the microstructure evolution in the affected layer (Deng et al., 2023).

The scanning pattern has a marked effect on the efficacy of LSP on fatigue life enhancement. FEM simulations and experimental results by Adu-Gyamfi et al., (2018) have shown that the CRS layer is strongly affected by the LSP scanning patterns in aluminium alloy AA2024 (Adu-Gyamfi, et al. 2018). Three scanning patterns were investigated. While all scanning patterns showed an increase in fatigue life, the highest increase in fatigue life experienced with an L-spiral scanning pattern (Adu-Gyamfi, et al. 2018). While in LSP of AA7075 notched samples in bending fatigue (Peyre et al., 1996), it has been shown that there is a dramatic improvement in the fatigue life with clear differences in the early and later stages of crack growth, where the fatigue life improvements from LSP can be separated into a seven-fold increase in the early crack growth

stage and only a three-fold increase in the later propagation stage. For AISI 316L stainless steel, Correa et al. (2015) found that a fatigue enhancement from LSP can be increased from 166% to 471% by optimising the pulse sequence and scanning pattern, and that greatest improvements were seen when the crack propagation direction is perpendicular to the fatigue load (Correa et al., 2015).

The aim of this work is to assess the effect of LSP on the fatigue life extension for three different high strength, high toughness materials, namely aluminium alloy AA7075, Ti-6Al-4V, and AISI316 stainless steel. These materials represent a variety of structures and characteristics, namely a face centred cubic (FCC) material with a high stacking fault energy (SFE), a material with an HCP structure, and an FCC material with a low SFE, respectively. These characteristics will influence the efficacy of the LSP in each case. Therefore, the penetration depth of the LSP induced case hardened layer and its effect on the crack initiation zone are quantified for each material. Through this, the relative material specific efficacy of LSP for fatigue life improvement can be evaluated.

Experimental procedure

Selected materials and fatigue testing set-up

The three materials selected for this investigation were AA7075 aluminium alloy in the T6 condition, Ti-6Al-4V titanium alloy, and AISI 316 stainless steel alloy. The material was sourced in extruded rod form with varied diameters. To standardise the sample geometry, surface finish and properties, all samples were turned on a lathe to a final diameter of 12 mm and were partitioned to 80 mm in length. Circular cross section samples were chosen to eliminate the effects of laser pulse area overlap (Zhao et al.,2017). During bending fatigue, the specimen span was equivalent to four times the diameter (i.e., S = 4D), which corresponds to dimensions suggested in both the ASTM E1290-02 (ASTM Standard E1290-02, 2002) and E399-90 standards (ASTM Standard E399-90, 1990) for fracture toughness testing, although no Standard specifically makes use of un-notched, cylindrical bend test specimens for fatigue testing.

Laser shock peening and sample surface roughness

The laser shock peening process parameters are shown in Table 1. The laser shock peening parameters for each material were tested on typical Almen strips to determine the correct saturation intensity.

The non-LSP samples were polished for 10 seconds with 1200 grit paper, where the samples were rotated relative to the grinding media. The samples were polished to a surface finish corresponding to Ra of 0.2 micron prior to testing using diamond impregnated lubricant and cloth on a rotating sample. The surface roughness values were determined using a profilometer. The LSP process, performed on polished samples, introduces shallow, homogeneous dimples onto the surface of the specimen, related to the laser parameters. These dimples produced an Ra of 2.28 micron. To eliminate the effects of surface roughness on the fatigue life data,

Table 1

Laser peening parameters

Power intensity

Number of pulses

Wavelength

Water confinement

4.9 GW/cm2

800 spots per cm2

1064 nm

Spray nozzle

Fatigue performance improvement using laser shock peening

a number of the LSP samples were briefly polished after peening, to achieve a surface finish Ra of 0.2 micron. This small amount of polishing was assumed to contribute negligible surface hardening on the samples relative to the LSP process and was thus, not included as part of the case hardening investigation. AA7075-T6 samples were used to evaluate the effect of surface finish on fatigue life, and fatigue testing was performed on three conditions: (i) aspolished, non-LSP, (ii) LSP and, (iii) LSP-polished. For the titanium and stainless-steel samples, the tests were performed on as-polished non-LSP and LSP-polished (where the surface was polished to achieve a surface roughness of the order of 0.2 micron).

LSP penetration depth: hardness profiles and fracture surface analysis

The depth of penetration of the LSP and the CRS layer thickness was determined from a combination of hardness testing and fracture surface analysis. Incremental micro hardness measurements were taken at increasing surface depths on unfatigued specimens through successive polishing and indentations to obtain a hardness profile for increasing depth from the surface. Longitudinal sections of the test specimens, of a length L (between 15 mm and 20 mm) were mounted in resin with their surfaces parallel to their cylindrical axes exposed from the resin and incrementally polished, removing a surface layer of nominally 3 μm thick per polish step. A schematic of this process is shown in Figure 1. This process of polishing the surface of the specimen, measuring the thickness of the mount, and then making several HV indents on the surface, using the full length of the mounted specimen is used to increase the number of indents for each depth to improve statistical uncertainties in HV number. This increase in polishing depth is repeated until the total depth of polishing is approximately 1 mm, or where a sudden decrease in hardness was achieved consistently, indicating the end of the LSP penetration depth.

Determination of penetration depth through fracture surface analysis was performed through fractography on the fracture surfaces themselves and through crack path analysis on crosssections through samples after failure. The fractography analysis included the identification of ductile and brittle features, crack initiation sites (surface or sub-surface). For crack path analysis, samples representing a section through the fracture surface (i.e., perpendicular to the fracture surface) of the failed samples were cut, mounted, ground, and polished to a colloidal silica finish in preparation for viewing using light microscopy with a Nomarski prism. Rounding of the polished edges is due to the polishing process resulting in poor edge retention between the sample edge

and the mount resin, but this did not affect the interpretation of the crack path. Light micrographs allowed for the analysis of the crack path in the various materials and conditions. A deviation in the crack path is expected at a point corresponding to a sudden change in residual stress, such as the limit of the LSP penetration depth on the crack initiation side and the final fracture side of the samples.

Fatigue performance

The fatigue performance of the various test specimens was evaluated under cyclic loading using a 100 kN electro-servo hydraulic fatigue machine (ESH) at the University of Cape Town, Centre for Materials Engineering. Test specimens of all materials, in both the LSP and non-LSP conditions, were fatigued to failure using a 3-point bending system at a constant amplitude bending stress. The testing frequency was set at 10 Hz using a sinusoidal wave at a stress ratio (R) of 0.1.

For the AA7075-T6 material, the initial stress estimate used to calculate the required load for the fatiguing of the test specimens in the T6 material condition was determined by looking at stresses used to generate S-N curves for AA7075-T6 found in literature. An incremental bending fatigue approach was then implemented to adjust incrementally the stress level to a level at which crack initiation occurred at a desirable cycle count of approximately 30 000 cycles. Based on the superimposed curve, a fatigue life of 30 000 cycles could be achieved somewhere in the applied stress range of 482 MPa to 551 MPa, achieved through incremental increases in the fatigue testing stress level. A final stress of 585 MPa was found to result in specimens failing at approximately 30 000 cycles, and this stress was kept constant for the test specimens in their respective material conditions regardless of the surface treatment that they had received. The stress level of 1080 MPa was used for Ti-6Al-4V and 1150 MPa for AISI 316 stainless steel, both calculated using Equation 1.

Where d is the diameter of test specimen, L is the span length and σ is 90% of ultimate stress (σUTS). This consistent equivalent stress level allowed the effects of the varying surface treatments on the fatigue lives of the test specimens to be determined.

Microstructure observations and fractography

The cross section through the fracture surface of the failed samples was viewed to analyse the crack path in the different materials using light microscopy, where any changes in the crack path direction

Figure—1A schematic of the process used for hardness testing. The specimen is mounted as shown in (a) and a thin layer is polished off between each hardness testing sequence as indicated in (b), and then after an additional layer in (c)

Fatigue performance improvement using laser shock peening

in the region corresponding to the CRS and LSP penetration depth was evaluated. The study did not focus on crack initiation sites. Fractography was performed using an FEI NovaNano SEM scanning electron microscope (SEM) to examine the fracture surfaces of the samples in the LSP and non-LSP conditions. Characteristic features identified on the fracture surfaces during fractography were related to the failure modes and the influence of the laser peened layer of material at the surface.

Results and discussion

Laser shock peening penetration depth for all materials

The hardness results from the longitudinal sections of the test specimens were plotted. The schematic in Figure 1 describes the experimental approach for hardness mapping. In the AA7075-T6 condition, the surface hardness was 185 HV, with a drop-off to 167 HV after 1.2 mm depth. In the Ti-6Al-4V the hardness was 334 HV at the surface that extended to a depth of 0.2 mm, a subsequent decrease to approximately 320 HV until 0.95 mm, and a drop to 309 HV beyond 1 mm. For the AISI 316, the hardness gradually increased from 281 HV at the surface to 294 HV at a depth of 0.65 mm, and then rapidly decreased to a hardness of 275 HV at a depth of 1 mm. The normalised hardness is defined here as the change in hardness value (∆HV) relative to the measured hardness of the parent material, where the parent material hardness is the hardness unaffected by the LSP. The normalised hardness profiles are presented in Figure 2. The penetration depth for all three materials is of the order of 1 mm, and this result is consistent with results documented by Clauer (1997; Clauer 2001). The penetration depth would be expected to increase with an increasing number of passes of the laser peening, as was shown by Zhou et al. (2023) in the case of AA7075 over a range of laser process parameters and passes (ASTM Standard E399-90, 1990).

The fracture surface cross section of the failed samples was studied to identify the presence of crack path. Directional changes were attributable to a change in the properties of the material. This property change is directly related to the transition between

the LSP-affected material and the parent material. Figure 3 shows the cross section through a failed Ti-6Al-4V sample, while Figure 4 shows an equivalent AISI316 stainless steel sample. The image shows the fast fracture edge of the samples (as opposed to the crack initiation side) as the effect of a property change was more pronounced on the side of the sample. The non-perpendicular angle of the fracture surface to the surface edge is a result of the compression curl (Quinn et al. 2005). There is a noticeable deviation in the crack path for both the Ti-6Al-4V and the AISI 316 samples, as shown in Figure 3 and Figure 4, respectively. The crack path deviation typically corresponds to the change in properties at the transition of the case depth of the LSP. The measured depth at the step in the crack path is approximately 1 mm in both cases. This indicates a change in the residual stress of the material attributed to the residual stresses associated with the penetration depth of the LSP. These depths correspond to the depth measured with hardness results as illustrated in Figure 2. Similar depths of penetration for AA7075 are reported by Zhou et al. (2023), for Ti-6Al-4V by Dyer et al. (2024) and for stainless-steel by Wu et al. (Wu, et al., 2023).

Fatigue performance

Aluminium AA7075 – Effect of laser peening on fatigue life

The results of the fatigue testing for AA7075 in four different conditions can be seen in Figure 5. The four different conditions represent the effect of laser peening on polished AA7075-T6 samples and the effect of laser peening on unpolished samples. From these results we can determine the effect of laser peening on the fatigue life of AA7075 in the T6 aged condition.

The AA7075 samples in the non-laser shock peened condition (non-LSP) failed after an average of approximately 39 000 cycles, while the samples that had undergone laser shock peening achieved a maximum of approximately 211 000 cycles. This is an improvement of more than five times with the application of laser shock peening. These results are for samples with the same equivalent surface roughness value, Ra, of approximately 0,2. This was achieved by polishing the samples before and after laser shock peening.

Figure 2—Hardness trend for longitudinal sections for tested sample to show penetration depth of the laser peening. Hardness trends are represented by a normalised hardness value, determined relative to the base material hardness

Figure 3—Micrograph of the cross-section through the crack path of failed Ti-6Al-4V sample taken on the fast fracture edge for (a) non-LSP and (b) LSP, showing a change in the crack path at approximately 900 μm, corresponding to the penetration depth of the LSP

Fatigue performance improvement using laser shock peening

Table 2

Surface roughness values for AA7075-T6 samples and the number of cycles to failure for each condition

non-LSP / polished

AA7075: LSP / as peened

with 2.08 decrease in Ra AA7075: LSP / polished

AA7075: LSP / intermediate polished (0.74 Ra) 0.74

Aluminium AA7075 – Effect of surface roughness on fatigue life

The effect of surface roughness is determined by identifying the number of cycles to failure in the fatigue test results on laser shock peened and non-laser shock peened samples in the polished and unpolished conditions. The results are also shown in Figure 5. The surface roughness values are shown in Table 2. The table also shows the number of cycles to failure associated with each condition and the calculated improvement in fatigue life with the improved surface roughness owing to polishing.

For samples that have not been exposed to laser peening, there is an improvement of approximately 1.2 times in the fatigue life with a decrease in Ra from 0.74 to 0.22. For samples that have been laser shock peened, in the as-peened condition there is a surface roughness of 2.28, which results in a fatigue life of 102 700 cycles. After peening followed by surface polishing to improve the Ra to 0.2, there is an improvement of the fatigue life in the order of about twice the fatigue life. This shows that surface roughness does play a very important part in the fatigue life, but also shows that the change in surface roughness does not fully account for the improved fatigue life after laser shock peening, where, for the

equivalent roughness values, the un-peened material exhibits an 18% improvement, while the peened exhibits a 44% improvement.

Titanium alloy Ti-6Al-4V

The fatigue life data for Ti-6Al-4V can be found in Figure 6. The fatigue life for Ti-6Al-4V with no laser peening was approximately 27 000 cycles, while after laser peening the fatigue life improved by more than five times, to approximately 143 300 cycles. This shows that titanium alloy Ti-6Al-4V responds very well to laser peening in the context of fatigue life extension, with a 5.3 times larger fatigue (81%). This is a far greater increase than the 22.2% increase documented by Zhang et al . (2010). The discrepancy between this work and Zhang may be due to a number of contributing factors, namely, the surface roughness, which is not explicitly stated in the work by Zhang et al. (2010), the testing parameters of the different tests, where the load fatigue force ratio was 0.3 as opposed to the 0.1 from this present work, or the underlying microstructure of the material, which had been specially heat treated to provide a bi-modal starting structure by Zhang et al. (2010). The difference in fatigue life increase owing to LSP may also be a result of a combination of the afore-mentioned effects.

Figure 4—Light micrograph of the edge side profile of failed AISI 316 stainless steel samples showing a change in the crack path at approximately 0.45 and 1.1 mm, corresponding to the penetration depth of the laser peening

Figure 5—Comparative table of fatigue results, for AA7075-T6 in LSP and non-LSP with a polished surface and unpolished surfaces

Fatigue performance improvement using laser shock peening

Stainless steel AISI 316

The fatigue life data for AISI 316 stainless steel can be found in Figure 6. The non-laser peened fatigue life was found to be approximately 32 400 cycles, while after laser peening this improved by 1.63 times to nearly 53 000 cycles. A similar investigation by Stamm et al. (1996) into the high cycle fatigue behaviour of austenitic stainless steel AISI 316L showed an increase in the fatigue life of 1.2 times. This is slightly lower than the 1.6 increase seen in this present investigation. This difference may be attributed to the fact that Stamm et al. (1996) did not polish the as-peened surface prior to fatigue testing. The increase in the fatigue life does indicate the relative effectiveness of laser shock peening on stainless steel. The reduced impact of LSP on fatigue life in FCC materials with low SFE has been highlighted by Deng et al., 2023 and is associated with dislocation mobility and twinning mechanisms in low SFE austenitic stainless steels.

Fractography

The fracture surfaces of the as-received and the laser shock peened tested samples were analysed and compared. For each tested sample and processing regime, images were captured in the crack initiation region; the crack propagation region within 1 mm of the edge of the specimen; the crack propagation region beyond the LSP case depth; and finally, within the fast fracture region.

The fractographic images revealed that the specimens subjected to LSP were indeed more brittle in the LSP zone than the non-LSP

specimens, as can be seen in Figure 7. This was indicated by brittle fracture characteristics, such as cleavage facets in Figure 7(b) for the non-LSP specimenAA7075-T6 and a smoother fracture surface. The spacing between fatigue striations on the fracture surfaces in the as-received specimens appear to be far greater than the spacing between fatigue striations in the LSP specimens, as seen in Figure 8(a) and Figure 8(b), respectively. The difference in fatigue striation spacing is in agreement with the results of Ren et al, (2013).

Measuring the distance between striations, it was found that the spacing in the as-received 7075-T6 is approximately 1.1 μm, while the spacing in the LSP 7075-T6 is approximately 0.4 μm. The smaller spacing implies a more brittle material that develops because of compressive residual stress and the reduced dislocation mobility owing to LSP (Ren et al., 2013). A similar trend can be seen in Figure 9 for AISI 316 where, by performing an approximate measurement of the distance between fatigue striations on the micrograph, it was found that the measured spacing between striations in the as-received specimen is approximately 0.7 μm, whereas the measured spacing between the striations in the LSP specimen is approximately 0.3 μm. Thus, the specimen that has been subjected to LSP is more resistant to crack growth than the as-received specimen, indicating that there is greater residual stress in the LSP zone.

The fractography in Figure 10 (a) shows an AA7075-T6 sample where the crack initiation site is sub-surface, at 0.3 mm away from the surface (the annotations A, C, and D represent crack initiation

Figure 6—Comparative fatigue results for all materials tested, in LSP and non-LSP with a polished surface

Figure 7—Crack initiation area of (a) an un-LSP AA7075-T6 sample, showing ductile tearing, and (b) a LSP AA7075-T6 sample, showing brittle cleavage facets

Figure 8—Fatigue striations within the LSP penetration zone on the fracture surface of AA7075-T6 in (a) un-LSP and (b) LSP samples. (Both images at same scale, as per micron marker in (b))

Fatigue performance improvement using laser shock peening

site, cleavage and dimpled fracture surface (ductile tearing), respectively). The applied bending stress combines with the peening induced residual stress, creating a resultant stress profile, where the highest tensile stress on the profile is just below the surface hardened layer, as shown in Figure 10 (b). The schematic diagram in Figure 10 (b) illustrates the combined residual LSP stress and pure bending stress, two triangles across the bending section representing tension on the side opposite the load point of 3-point bending, and compression on the opposite face. The schematic also indicates the presence of the residual stress from peening, which is compressive but is limited to a narrow subsurface depth. The LSP CRS layer tapers away deeper, but still within a confined layer close to the surface, into the material towards the neutral axis, where the tensile bending stress begins to dominate. The resultant of the two stresses reduces the surface tensile stress substantially and puts the peak tensile stress at some point inside the surface from which the crack initiates and radiates. After some fatigue loading, there is no evidence of a surface crack, as it is growing radially from this highest stress regime sub-surface. Suddenly the crack breaks through the surface and at this point, the sample goes from having no noticeable surface crack to having a visible surface crack driven by the growth of an internal crack of between 5 mm to 10 mm long, showing that it grew from the inside radially in all directions and finally broke through the surface.

Figure 11 (b) shows that, in the aluminium LSP specimens, crack initiation occurred just beneath the surface of the specimen, as opposed to at the surface in the non-LSP sample in Figure 11 (a), where the crack initiation point is clearly at the surface and would be the result of typical persistent slip banding as part of the initiation process. In Figure 11 (b), the chevron marks point to the crack initiating at the edge of the transition area where there is a change in the residual stress in the material, which corresponds to approximately 0.8 mm. The expansion of the initiation region after LSP indicates that for samples that have been exposed to laser shock peening, many fatigue cycles are spent on the crack initiation phase before crack propagation occurs in a steady state manner. Figure 6 shows features on the fracture surface within the

crack initiation region for AA7075-T6 and highlights the change in the material because of the LSP process. The LSP sample exhibits cleavage facets on the fracture surface, which are features associated with brittle materials, while the non-LSP sample exhibits dimples, which are typical ductile features. This same trend was seen in both the Ti-6Al-4V and the AISI 316. However, in the titanium and stainless steel the crack initiation region was smaller than that of the aluminium alloy. The change in depth of the transition region is also illustrated in the hardness profiles in Figure 2. This indicates that materials have different susceptibilities to the effects of the LSP process. The susceptibility is related to the crystal structure and the SFE of the material, which dictates the slip planes available for deformation and the mobility of the dislocations under load.

Conclusions

Laser shock peening (LSP) significantly enhances the fatigue performance of various materials. The efficacy of LSP is approximated by the relative increase in fatigue life for the peened versus non-peened conditions. Three materials were considered, aluminium AA7075-T6, an FCC material with a high SFE, AISI 316 austenitic stainless steel, an FCC material with a low SFE and Ti-6Al-4V, and an HCP and BCC two-phase structure. All three materials are considered ductile, high strength materials with good corrosion resistance properties. The results indicate that aluminium AA7075-T6 with an FCC structure and a high SFE responds best to LSP and shows the greatest increase in fatigue life in the LSP condition.

The main findings of the work are summarised below:

➤ In the case of aluminium alloy AA7075-T6, LSP extended fatigue life by more than five times when the surface roughness was polished to an R_a of 0.2 microns, matching that of the non-LSP sample.

➤ For as-LSP treated AA7075-T6, the surface roughness was measured at 2.28 microns, and this untreated roughness still resulted in a 4.1-fold increase in fatigue life. When the surface roughness was equalised at 0.2 microns, LSP improved the fatigue life of AA7075-T6 by 5.4 times. This indicates that

Figure 9—Fatigue striations within the LSP penetration zone on the fracture surface of AISI 316 stainless steel in (a) un-LSP and (b) LSP samples. (Both images at same scale, as per micron marker in (b))

Figure 10—(a) AA7075-T6 fracture surface showing the crack initiation point sub-surface at approximately 0.25mm. This is the location where the maximum resultant stress would be located, as indicated in the schematic in (b) [32]

Fatigue performance improvement using laser shock peening

11—AA7075-T6

reducing the surface roughness—without compromising the LSP-induced case depth—is key to maximizing fatigue life across different materials.

➤ For titanium alloy Ti-6Al-4V and AISI 316 stainless steel, LSP increased fatigue life by 5 times and 1.7 times, respectively. This demonstrates that titanium alloy Ti-6Al4V and aluminium alloy AA7075-T6, both ductile at room temperature, are far more responsive to LSP than AISI 316 stainless steel.

➤ Results indicate that the improvement in fatigue life can be attributed to the residual compressive stresses induced by LSP (although not quantified in this work), rather than the penetration depth, which was approximately 1 mm for all three materials tested.

➤ LSP samples exhibited a notable expansion in the crack initiation zone for all materials and a sub-surface crack initiation point, as evidenced by the features on the fracture surfaces.

Future work in this area must include quantification of the compressive residual stress in each material and the dislocation structures in the CRS layer that are linked to crystal structure and dislocation mobility.

In conclusion, LSP is an effective method for improving the fatigue life of components made from aluminium AA7075-T6, titanium Ti-6Al-4V, and AISI 316 stainless steel. However, the degree of improvement varies between materials, with some being more susceptible to the benefits of LSP-induced compressive residual stress, which penetrates up to 1 mm into the material.

Acknowledgments

Acknowledgments to Daniel Glaser at the National Laser Centre (NLC) at the Council for Scientific and Industrial Research (CSIR) in South Africa for the laser peening of the investigated samples, and to the Electron Microscope Unit (EMU) at the University of Cape Town for the use of the electron microscopes and operator guidance by Miranda Waldron.

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Figure

fracture surface (a) un-LSP and (b) LSP, with the extended crack initiation area owing to the residual stress from LSP, where a sub-surface crack would have initiated

Affiliation:

1Department of Geography and Resource Development, University of Ghana, Accra, Ghana

2Institute of Earth and Space Sciences, Eskisehir Technical University, Eskisehir, Turkey

Correspondence to:

C. Kwang

Email: ckwang@ug.edu.gh

Dates:

Received: 30 Sept. 2021

Revised: 3 May 2025

Accepted: 15 May 2025

Published: July 2025

How to cite:

Kwang, C., Uygycgil, H. 2025. Comparison of remote sensing techniques in mapping hydrothermal alteration associated with Ovacik epithermal gold-silver mineralization. Journal of the Southern African Institute of Mining and Metallurgy, vol. 125, no. 7, pp. 355–360

DOI ID:

https://doi.org/10.17159/2411-9717/1775/2025

ORCiD:

C. Kwang

http://orcid.org/0000-0003-3545-1429

H. Uygucgil

http://orcid.org/0000-0003-3100-0129

Comparison of remote sensing techniques in mapping hydrothermal alteration associated with Ovacik epithermal gold-silver mineralization

by C. Kwang1, H. Uygycgil2

Abstract

The use of remote sensing and geographic information system techniques in mineral exploration is still underutilised. There is a need to explore the potential of new robust remote sensing and geographic information system techniques in the field of mineral exploration. In this study, we mapped the argillic alteration associated with epithermal gold-silver mineralisation. A band ratio of band5/band6 and Crósta technique on Advanced Spaceborne Thermal Emission and Reflection Radiometer within the Google Earth Engine Environment was applied. The bands 5 and 6 were selected for the band ratio after critically analysing and comparing the illite spectral extracted from the Advanced Spaceborne Thermal Emission and Reflection Radiometer image and the illite spectral obtained from the United States Geological Survey spectral library. The band ratio and Crósta technique identified and mapped the Ovacik goldmine and other argillic alteration zones. Even though both methods mapped out the argillic alteration zones associated with gold-silver mineralisation, the Crósta technique produced the optimal argillic alteration. The study demonstrates the analytical capabilities of Google Earth Engine in image processing and analysis for mineral exploration, such as discovering hydrothermal alteration associated with mineralisation.

Keywords band ratio, Crósta technique, argillic alteration, epithermal gold-silver deposit

Introduction

Minerals and metals are vital for the sustainability of humanity on Earth and are used in innumerable and diverse applications, for instance, from building aircraft to producing ornaments for human use. Economically, mineral resources contribute to gross domestic product and offer employment opportunities. According to Creamer (2012), mining's contribution to the world's gross domestic product is approximately 45 per cent. Thus, sustainable mining of minerals and metals remains paramount. in most instances, mining usually only commences after extensive mineral exploration.

Despite the benefits, mineral exploration is one of the highest risk ventures in recent times in terms of the associated cost and failure (Gandhi, Sarkar, 2016). Modern technologies, such as remote sensing and geographic information systems (GIS), have been used as an alternative approach for mineral prospecting to reduce the cost and failure of new minerals' discovery. Remote sensing and GIS have yielded many successful mineral exploration works (Ninomiya et al., 2006; Chen et al.; 2007; Tangestani et al., 2011; Hede et al., 2015; Salehi et al., 2017; Rajendran, Nasir, 2017; Wang et al., 2017). Therefore, the objective of the study was to use Google Earth Engine as a working environment for the identification and mapping of the argillic alteration zones associated with epithermal gold-silver deposits in Izmir, Turkey.

Geological setting of the study area

The Ovacik gold mine is located in the northern part of Izmir in Turkey, lying adjacent to the eastnortheast (ENE) trending Bergama graben (Figure 1).

The first economic epithermal deposit found in Turkey using modern exploration technology is the Ovacik gold-silver deposit (Yilmaz, 2002). The principal rock types within the Ovacik mine and its environment include limestones, palaeozoic metamorphic rocks, and intrusive andesitic rocks (Yilmaz, 2002), as illustrated by Figure 2. The main minerals associated with the argillic alteration of the Ovacik epithermal deposit are illite, interlayered illite/smectite, illite/chlorite, adularia, smectite, chlorite, and chalcedonic quartz (Yilmaz, 2002). This study area also confirms illite minerals as one of the main argillic alteration minerals in Ovacik.

Comparison of remote sensing techniques in mapping hydrothermal alteration

Material and methods

The materials used in this study include an Advanced Spaceborne Thermal Emission and Reflection Radiometer (ASTER) image, a geological map of the Ovacik Mine, and Google Earth Engine Application Programming Interface (API). In order to identify and map the argillic alteration associated with gold-silver mineralisation within the study area, Google Earth Engine API was used to extract the image spectra, perform the band ratio, and the Crósta technique was applied to the image. Google Earth Engine has reduced the burden of downloading satellite images before performing image processing and analysis. Google Earth Engine can process broad coverage areas image analysis faster and easier than most commercial remote sensing and GIS software (Gorelick et al., 2017).

The Ovacik gold mine was selected because its characteristics are similar to the other gold-silver mineralisation locations in Turkey. The ASTER image captured on 23 September 2006, was first loaded on the Google Earth Engine environment. Owing to its diagnostic characteristics and predisposition to clay minerals detection and mapping, shortwave infrared bands were selected. In brief, the image was transformed from radiance to reflectance using gain and offset values of bands of the image before the band ratio operation was performed. The geological map of the Ovacik gold mine was also imported to Google Earth Engine, and sample points were extracted from known mineralised locations. Sample points extracted from the known mineralised locations in the geological map were used to obtain image spectra from the ASTER image. The illite mineral spectra obtained from the United States Geological

Figure 1—Google Earth Image of the Ovacik gold mine

Figure 2—Geological map of the Ovacik area. Source: (MTA, 2018)

Comparison of remote sensing techniques in mapping hydrothermal alteration

Band 7 Reflectance 0 0.5 1 1.5 2 2.5 3 3.5 Wavelength (mm)

Illite (USGS spectral library)

Point 1 (Extracted ASTER image spectral)

Point 2 (Extracted ASTER image spectral)

Illite (USGS spectral library)

Diagnostic AlOH absorption ASTER Bands

Point 1 (Extracted ASTER image spectral)

Point 2 (Extracted ASTER image spectral)

(a) (b)

Survey (USGS) spectral library was compared with that of the spectra acquired from the ASTER image, as illustrated in Figure 3. The extracted image spectrum shows the diagnostic characteristics of the illite mineral, which is a known mineral of argillic alteration. In Figure 3, the illite mineral showed high reflectance in Band 5 and strong absorption, or low reflectance in Band 6, which were vital diagnostic characteristics for its identification and mapping using the band ratio method.The presence of argillic alteration characterised by illite mineralisation was also identified by performing the Crósta technique using bands 1, 3, 5, and 6, as suggested by Crósta et al. (2003). All the processed images were exported from the Google Earth Engine as Geographic Tagged Image File Format (GeoTIFF) images for map creation.

Band ratio

Minerals and rocks do not show similar reflectance responses at different ranges of spectral wavelengths (Chandrasekar et al., 2011). Band ratio operation can be easily performed to map hydrothermal alteration by using the spectral diagnostic characteristics exhibited by rock and minerals. For instance, in 800–1000 nm, iron oxide minerals portray high absorption features and high reflectance in other wavelength regions (Pontual et al., 2008). In mapping hydrothermal alteration zones associated with mineralisation, the most common techniques applied are the band ratio and principal component analysis (PCA) (Yajima, 2014). There is a reduction in the effect of the seasonal changes, slope shadows, and the elimination of sunlight angles effect when band ratio analysis is applied on satellite imagery (Jensen, 1986).

Crósta technique

The Crósta technique is a modified version of the standard PCA, and it utilises four image bands (Loughlin, 1991). The Crósta technique uses the linear combinations of the PCA to determine which principal component (PC) image best describes the theoretical spectral response of the mineral under investigation (Loughlin, 1991). The Crósta technique has the advantage of determining whether the target feature type has been enhanced in the previous principal component image as dark or bright pixels. Other image analyses, such as band colour combination and band ratio, were performed to decide the band that could best show the

argillic alteration within the study area before choosing the bands for the Crósta technique.

Results and discussion

The short-wave infrared (SWIR) bands offer a moderate spatial and spectral resolution for detecting and identifying clay and hydrous minerals. Within the SWIR bands, the clay and hydrous minerals reveal diagnostic spectral characteristics. In the SWIR region of the electromagnetic spectrum, illite minerals portray different absorption features. The wavelengths of distinct absorption features identified within the SWIR include 1.900 µm (water absorption), 2.180-2.228 µm (AlOH), 2.342 µm, and 2.435 µm (Pontual et al. 2008). Field studies of the hydrothermal alteration of the Ovacik gold deposit reveal the abundance of illite minerals and its association with epithermal gold-silver mineralisation. The image spectra, as shown in Figure 3, revealed the diagnostic feature of illite mineral as 2.180-2.228µm (AIOH) after being critically examined. At a wavelength of 2.185-2.225µm, the AI-OH molecular vibration process obtained the high AIOH absorption feature (Pontual et al., 2008) corresponding to band 6 of the ASTER image.

Table1 shows eigenvector loadings, eigenvalues, the percentage variances of individual principal component images of the bands used in the Crósta technique, and statistics of individual bands used to calculate the eigenvector of the Crósta technique. Each of the individual principal components (PC) in Table1

Table 1

The eigenvector loadings, eigenvalues, and percentage variance of bands (1, 3, 5, and 6)

Figure 3—The extracted ASTER image spectra within the (a) VNIR-SWIR and (b) SWIR region and the USGS illite spectra

Comparison of remote sensing techniques in mapping hydrothermal alteration

has a corresponding principal component image. The principal components in Table1 indicate the numeric values of argillic minerals. Principal component images display spectral information of the argillic minerals as grey in colour. Principal components, which contained the argillic alteration characterised by illite mineralisation were based on Loughlin's (1991) criterion, which states that "PC that contains the target spectral information shows the highest eigenvector loadings from the image bands, coinciding with the target's most diagnostic features, but with opposite signs (+ or -)". Based on Loughlin's criterion, PC3 was selected as it has the highest value with high opposite loadings values. The PC3 has

a high eigenvector loading value for bands 5 (0.4) and 6 (-0.76). The high and negative eigenvector loading of band 6 and the low positive value of band 5 are the main reasons why argillic alteration zones in Figure 4 appeared as dark rather than bright pixels. In order to display the argillic alteration zones as bright pixels, the spectral values of PC3 were multiplied by negative 1 (Tangestani, Moore, 2000). Figure 3 shows high absorption of illite in band 6 and high reflectance in band 5.

A threshold value of 1.08, calculated by using the mean plus twice the standard deviation, was applied at a 95% significance level on the image processed with the Crósta technique method to

Figure 4—PC3 image of Crósta Technique for argillic alteration zones detection

Figure 5—Comparison of the Argillic alteration zones from band ratio (b5/b6) and Crósta technique (1, 3, 5, and 6)

Comparison of remote sensing techniques in mapping hydrothermal alteration

separate altered from non-altered argillic zones.The orange colour in Figure 5 shows zones with reflectance values above the threshold value, and these are the most likely argillic zones associated with gold-silver mineralisation. When the argillic alteration zones extracted from the band ratio of band 5/band 6 were overlaid with alteration zones produced from the Crósta technique, most of the argillic alteration zones from band ratio did not coincide with alteration zones of the Crósta technique. The argillic alteration zones obtained from the band ratio were affected by vegetation cover, while in the Crósta technique, the effect of vegetation was suppressed. Suppression of the vegetation effect helped enhance spectral reflectance of the illite mineral.

The spectral characteristics of minerals are not vastly distinctive, and this must be considered when mapping hydrothermal alteration. Although the wavelength for reflectance bands may vary, many clay minerals' diagnostic absorption features fall within almost the same wavelengths. This suggests that the same band ratio index may not produce optimal hydrothermal alteration zones in all geographic locations. Band combinations can aid in selecting suitable bands for hydrothermal alteration identification and mapping. By comparing the argillic alteration zones produced from two methods, the Crósta technique produced better alteration results than the band ratio method. However, both the band ratio and Crósta technique identified the Ovacik goldmine within the argillic alteration zones.

Conclusion

This study demonstrates the use of Google Earth Engine (GEE) for detecting and mapping argillic alteration zones associated with gold-silver mineralisation in Izmir using the band ratio and Crósta techniques. Despite its potential for handling large geospatial datasets efficiently, the use of GEE in mineral exploration remains limited.

The study evaluated methods for identifying and mapping argillic alteration zones linked to epithermal gold-silver mineralisation. Specifically, it assessed the performance of Landsat bands 5 and 6 for band ratio analysis and bands 1, 3, 5, and 6 for the Crósta technique. The findings indicate that the Crósta technique outperformed the band ratio method, as it leverages spectral bands associated with alteration minerals, providing greater accuracy and reliability in delineating areas with potential gold-silver deposits.

Furthermore, this study underscores the transformative potential of Google Earth Engine as a powerful geospatial tool for advancing mineral exploration. Its ability to process and analyse large-scale satellite imagery enhances the efficiency and precision of remote sensing-based mineral prospecting, offering a valuable approach for future exploration initiatives.

References

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SAIMM PYROMETALLURGY

INTERNATIONAL CONFERENCE 2026

Foundations of Competitiveness and Sustainability

Comparison of remote sensing techniques in mapping hydrothermal alteration

INTRODUCTION

The South African metallurgical industry is facing significant obstacles: Rising production costs, the closure of key service providers like refractory suppliers, and increasing pressure globally for the industry to transform. To face these challenges, it is important for different areas of the industry to work together.

The Southern African Institute of Mining and Metallurgy (SAIMM) invites you to the 2026 Pyrometallurgy International Conference, taking place from 25 to 29 May 2026. The conference will bring together professionals and experts from the fields of pyrometallurgy, furnace tapping, and refractories under the theme: Foundations of Competitiveness and Sustainability. This theme challenges us to collaborate in finding answers to the difficult questions we are faced with and to develop sustainable solutions to safeguard the future of our industry. We are confident that through collaborative research, innovative technology development and operational excellence, we can transform the industry and secure a greener future for the next generation of pyrometallurgists.

By integrating these three key focus areas, we aim to create a platform for strategic thinkers, policymakers, researchers, and economic influencers to share insights, challenge assumptions, and collaborate on practical solutions for the future of the industry.

WHO SHOULD ATTEND

• Professionals and researchers focused on pyrometallurgy, furnace tapping and refractories

• Industry leaders

• Academics

• Students

25 May 2026 – Workshops

26-28 May 2026 – Conference

29 May 2026 – Technical Visits

DAY 1: WORKSHOPS

• The Future of Pyrometallurgy in South Africa

• Advanced Tapping and Refractories

DAY 2-4: SYMPOSIA (PARALLEL)

• Pyrometallurgy in SA Under Pressure – What Next?

• Tapped In – The Future of Sustainable Furnace Tapping

• Quo Vadis, Refractories?

DAY 5: TECHNICAL SITE VISITS

• To Be Confirmed

CALL FOR PAPERS, PRESENTATIONS AND POSTERS

Papers, presentations or posters are invited on any topic related to the conference and can be submitted to any of the three Symposia.

Prospective authors are invited to submit titles and abstracts of their papers in English. The abstracts should be no longer than 500 words and submitted via the SAIMM Abstract Portal.

Acceptance of papers for publication in the SAIMM Journal will be subject to peer review by the Conference Committee and SAIMM Publications Committee pre-conference.

ABSTRACT

KEY DATES

• Abstract submission deadline: 29 September 2025

• Paper submission deadline: 10 November 2025

• Workshops: 25 May 2026

• Conference: 26-28 May 2026

• Technical Visits: 29 May 2026

Affiliation:

1Department of Materials Science and Metallurgical Engineering, University of Pretoria, South Africa

Correspondence to:

A.M. Garbers-Craig

Email:

Andrie.Garbers-Craig@up.ac.za

Dates:

Received: 20 Dec. 2024

Revised: 18 Feb. 2025

Accepted: 9 May 2025

Published: July 2025

How to cite:

Garbers-Craig, A.M., Cameron, I., Ramjee. S. 2025. Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch as binder. Journal of the Southern African Institute of Mining and Metallurgy, vol. 125, no. 7 pp. 361–370

DOI ID:

https://doi.org/10.17159/2411-9717/3623/2025

ORCiD:

A.M. Garbers-Craig

http://orcid.org/0000-0002-0298-8097

I. Cameron

http://orcid.org/0000-0002-6065-3690

S. Ramjee

http://orcid.org/0000-0003-0104-207X

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch as binder

by A.M. Garbers-Craig1, I. Cameron1, S. Ramjee1

Abstract

The cause of reduced workability and increased ageing of a blast furnace taphole clay was examined. The investigated taphole clay contained 60 mass% alumina, with phenol-formaldehyde resole resin and liquid pitch serving as the binder system. The workability and Marshall extrusion pressure of the as-manufactured clay samples were evaluated to assess the extent of ageing of the clay. The wettability of all the raw materials was investigated to confirm compatibility between the dry raw materials and both the resin and liquid pitch, while the dry raw materials were analysed using XRF, XRD, and SEM-EDS. The characterisation of the resin and liquid pitch, as well as an analysis of their interaction, was performed using viscosity measurements, Fouriertransform infrared spectroscopy (FTIR), thermogravimetric analysis (TGA), and differential scanning calorimetry (DSC). The analyses confirmed a chemical interaction between the resole resin and liquid pitch, where the chemical structure of the resole resin changed when mixed with pitch, preventing the typical curing behaviour of the resin. Upon ageing of the resin-pitch mixtures, the resin underwent premature cross-linking, causing the curing process to initiate at lower temperatures, i.e., earlier onset of curing. This premature cross-linking was likely due to secondary amines present in the liquid pitch, which acted as a catalyst for the curing process of the resin. This reduction in curing temperature after ageing was confirmed by an increase in binder viscosity, which was identified as the primary cause of reduced workability, accelerated ageing, and increased Marshall extrusion pressure of the taphole clay.

Keywords taphole clay, phenol-formaldehyde resin, liquid pitch, workability, Marshall extrusion pressure

Introduction

Blast furnace taphole clay has been used for decades with little change to the primitive raw materials and binders it employs. Recent developments have explored the use of cheaper alternative raw materials as aggregates to reduce the cost of the clay, as well as tar-alternative binders that are less harmful to both the environment and operators who manufacture or use the material (Rosch, 2019). A key objective of these developments was the creation of a more environmentally friendly “green” taphole clay, which led to the present investigation.

The change in the binder introduced in this product development involved replacing the conventional high-volatile tar and resole resin combination with a synthetic liquid pitch and resole resin. However, a problem was identified with this new binder system: premature ageing of the clay, which manifested as a crumbling effect and reduced workability. The objective of this investigation was therefore to determine why this new binder combination caused the clay to age prematurely.

Several factors influence the ageing and workability of taphole clay, including the particle size distribution (PSD) of individual raw materials; the solubility of aggregate and matrix raw materials in the binder or binder combination; impurities in the aggregates and matrix (such as lime and sulphur); ageing temperature; and the viscosity of the binder, which is in itself influenced by temperature and the degree of cross-linking (Nelson, Hundermark, 2014).

The PSD of both the individual raw materials and the overall mixture affects the rheological behaviour of the clay. Any rheological change, in turn, impacts the workability of the taphole clay (Adeyinka et al., 2009). The solubility of the aggregate and matrix raw materials in the binder or binder mixture also influences workability through polar-polar or nonpolar-nonpolar interactions between the solids and the binder(s). The polarity interactions, whether liquid-liquid or liquid-solid, determine

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

whether the clay remains well mixed or separates upon ageing (Lin et al., 1993). Mismatched polarities between the binder components can lead to reduced workability and increased ageing.

The viscosity of the binder or binder combination is affected by both temperature and structural changes caused by cross-linking. As the binder undergoes structural changes during cross-linking, its viscosity increases, leading to a reduction in workability (Li et al., 2018; Moritz, 1989). Premature cross-linking of the resin in the binder accelerates the curing process (Laza et al., 2001), causing a premature increase in binder viscosity, which reduces workability and increases ageing of the taphole clay. Certain impurities, such as lime or sulphur, associated with the raw materials, can also catalyse cross-linking of the resin in the binder (Young, Lovell, 1991).

During the synthesis of resole resin, phenol reacts with formaldehyde under basic conditions to form an addition compound (due to chemical addition), followed by a condensation reaction. This condensation step produces methylene bridges with methylol functional groups in ortho or para positions (Laza et al., 2001). Due to the water present in the resole resin, various types of catalysts can be used for curing, including acids, bases, or amines (primary, secondary, and tertiary) (Laza et al., 2001). Hexamethylenetetramine (hexamine) is often used as a curing agent in resole-type resins.

Experimental procedure

This study was divided into three parts: In the first part, the properties of two taphole clays with different binder systems were studied. The second part focused on the aggregate and matrix raw material used to constitute the taphole clay, while the third part investigated the individual binders as well as different binder combinations. Tests performed on the taphole clays included workability, Marshall extrusion pressure, and viscosity. The aggregate and matrix materials were analysed for particle size distribution (PSD), polarity, wettability, and associated impurity components, while the binders and binder combinations were evaluated based on their structural and thermal characteristics.

Materials

The two taphole clay samples (Clays A and B) were prepared according to the formulation shown in Table 1 and in accordance

with ASTM C1054-18 (2018) (Saeki, Tanaka, 1984). The amounts and size distributions of the aggregate and matrix materials used in taphole Clays A and B were similar, the only difference lay in their binder systems. The particle size distributions of Clays A and B were continuous, with no gaps that could negatively affect the flow properties of the clay when extruded or pushed through a mud gun. Taphole Clay A contained 17% liquid pitch and 1% resin, while Clay B only contained 18% liquid pitch and no resin. The aggregate and matrix materials accounted for 100% of the formulation by mass, with the resole resin and liquid pitch added on top of that amount. Hence, the designation “+100%” in Table 1.

The aggregate included bauxite and andalusite, while the matrix material comprised andalusite (ball mill fines), calcined clay (ball mill fines), and calcined alumina. Kaolinite was added as filler and sintering aid, coal was included to enhance gas permeability, and silicon carbide was added to improve abrasion resistance.

The binders used were resole resin and liquid pitch, as well as various combinations of the two binders (5.6% resin – 94.4% pitch, 25% resin – 75% pitch, 50% resin – 50% pitch, 75% resin – 25% pitch).

A defined protocol was followed for the ageing and curing tests. Binder samples were prepared by weighing the appropriate ratios of liquid pitch to resole resin and mixing them for 5 minutes. After mixing, the viscosities of the samples were measured, and the samples were analysed using Fourier-transform infrared spectroscopy (FTIR) and differential scanning calorimetry (DSC). To simulate ageing, the samples were placed in an oven at 45 °C for two weeks. After this ageing period, the samples were re-analysed using the same techniques (viscometry, FTIR, and DSC). For the curing tests, the same protocol was followed, except that samples were cured at 130 °C for 4 hours to ensure complete curing of the resin.

Test methods used to study the taphole clay

The workability and Marshall extrusion pressure (MEP) of a taphole clay are closely related and therefore cannot be considered in isolation. Workability was assessed using the ASTM C181-11 (2018) test, which involves applying a load to a sample using a rammer and measuring the change in sample height. The MEP test was conducted using the configuration shown in Figure 1. A

Table 1

Taphole clay design used for both Clays A and B

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

450 g taphole clay sample was weighed and heated to 45 °C in a water bath for 2 hours. The heated sample was placed into the cylinder of the assembly, and the piston positioned on top of the sample. The assembly was inserted into the Marshall test automatic extruder, and the maximum force exerted on the sample was recorded. The MEP was then calculated as the average value obtained from Equations 1 and 2. The constants used in both Equations were calculated based on the setup shown in Figure 1. The maximum pressure on the piston/die was expressed in MPa, where the maximum load was measured in kN and the curved surface area in mm2.

Test methods used to characterise the aggregate and matrix materials

Wettability

The wettability of the aggregate and matrix materials with respect to polar (water) and non-polar (hexane) liquids was examined to confirm that there is no polarity mismatch between solid particles (i.e., fines and aggregate) and the binders (i.e., resin and liquid pitch). The polarity match between binders and particulates is important for two reasons. Firstly, the solubility of the resin in the pitch is highly dependent on the polarity of the compounds it contains and its molecular mass distribution (Li et al., 2000). A polarity mismatch may cause problems with the mutual solubility of the resin and pitch, as well as their interaction with the particulates in the clay (Lin et al., 1993). Secondly, a polarity mismatch between one of the particulate raw materials and the rest of the oxides, carbides, and binders may cause the clay to separate due to repulsive forces within the clay mixture. This separation can negatively affect the clay’s workability over time.

The phenolic resin is considered polar because of its functional groups, but this is dependent on the ratio of phenol to formaldehyde (Li et al., 2000). The liquid pitch is also considered to be polar due to one of its major constituents being coal tar pitch. The coal tar pitch has hydroxyl and amino functional groups, which convey more polar behaviour to this liquid (Knicker et al., 1996; Kabe et al., 2004).

Wettability was evaluated using an Attention Tensiometer Sigma 700/701 in conjunction with Washburn’s equation, to determine contact angles and wettability (Galet et al., 2010). During the wettability test, the sample was suspended in a test medium: water for the polar wettability test and hexane for the non-polar wettability test. The tensiometer recorded the change in sample mass over time. A mass2/time value was obtained from the slope of the mass versus time graph. Washburn’s equation relates this mass2/ time value to the contact angle:

Where Cw is a constant calculated for each sample, ρl the density of the test liquid in kg/m3, μ the viscosity of the test liquid in Pa.sec, γ being the surface tension of the test liquid, and θ the contact angle.

Mineralogical characterisation

Mineralogical characterisation was carried out to ensure the quality and to identify any impurities, such as lime or sulphur, that could attribute to cross-linking of the resole resin used in the binder. The mineralogical characterisation of the aggregates and matrix raw materials was done using X-ray fluorescence (XRF), X-ray diffraction (XRD) and scanning electron microscopy with energy dispersive spectroscopy (SEM-EDS) analyses. The crystalline phases present in the mineral aggregates were evaluated using XRD. Samples were pulverised and homogenised in accordance with the method described by Hanke (2017). The samples were loaded in accordance with the Panalytical backloading system and analysed using a PANalytical X’pert Pro powder diffractometer in θ-θ configuration with an Xcelerator detector and variable divergence and fixed receiving slits with Fe-filtered Co-Kα radiation (λ=1.789A) (Chatterjee, 2001). The mineralogy and microscopic properties of the aggregate and matrix raw materials were investigated using SEM-EDS. Backscatter electron images were captured and point and area analyses performed to supplement the XRF and XRD analyses. The quantitative EDS analysis was conducted in accordance with the ASTM e1508-12a (2019) standard guideline. The coal aggregate was characterised in terms of ash, moisture, volatile matter, fixed carbon, and sulphur contents in accordance with the ASTM D3172-13 (2013), SANS 927:2013 (2013), ASTM D5373-16 (2016), and ASTM D4239-18e1 (2018) test methods, respectively.

Test methods used to characterise the binders and binder mixture

Viscosity

The viscosities of Clay A, the resole resin, the pitch as well as resin-pitch mixtures were measured using an Anton Paar Physica MCR301 equipped with a Peltier heating chamber and 50 mm plate spindle. The rheology of taphole Clay A was studied by measuring its viscosity at different temperatures over time. Each binder was characterised by measuring the change in viscosity as a function of shear rate to distinguish between Newtonian and non-Newtonian behaviour. The resole resin and liquid pitch binders used in Clay A

Figure 1—Schematic of the Marshall extrusion mould assembly

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

were analysed to monitor cross-linking as a possible cause of ageing and reduction workability. Viscosity changes were also examined as a function of temperature, with samples heated from 30°C to 130°C at a heating rate of 10°C/min. Binder combination samples (5.6% resin – 94.4% pitch, 25% resin – 75% pitch, 50% resin – 50% pitch, 75% resin – 25% pitch) and aged binder combination samples (aged at 45°C for two weeks) were also tested for viscosity as a function of both shear rate and temperature over the same range.

FTIR/ATR

The resin, liquid pitch, and their combinations (as used in the viscosity measurements) were characterised using Fouriertransform infrared spectroscopy with attenuated total reflectance (FTIR/ATR). FTIR was used to identify the functional groups in the resin and pitch, and to detect any changes to these functional groups when the components were mixed and aged. FTIR/ATR analysis was employed to determine whether changes in chemical structure correlated with the observed viscosity variations in the mixed and aged samples as a function of temperature. Cured samples (130°C for 4 hours) were also analysed to track progressive changes in functional groups during mixing, ageing, and curing. These structural changes were then related to the observed viscosity trends. FTIR/ATR was conducted using a PerkinElmer spectrum 400 FT-IR, with samples placed directly onto the diamond cell for analysis. Background scans, as well as CO2 and H2O spectra, were removed from the sample spectra. Analyses were conducted at 25°C.

Differential scanning calorimetry

Differential scanning calorimetry (DSC) was used to investigate the heat flow in the binders and their combinations (mixed, aged, and cured) as a function of temperature. DSC analysis supported the viscosity and FTIR results by identifying characteristic endothermic and exothermic transitions in the resin, pitch, and their combinations. A small sample of binder (approximately 20 mg) was placed in a 30 μL aluminium pan, sealed with a lid, and heated in a nitrogen atmosphere. Samples were heated at 10°C/min in air, with a gas flow rate of 20 mL/min.

Thermogravimetric analysis

Thermogravimetric analysis (TGA) was performed to measure mass changes in the virgin binders as a function of temperature. A 20 mg sample of either resin or pitch was heated in an alumina crucible using a TA Q600 SDT. Samples were heated from 30°C to 130°C in a nitrogen atmosphere, and mass loss was recorded. The heating rate was 10ºC/min, with a nitrogen flow rate of 20 mL/min.

Results and discussion

Properties of Clays A and B

This section describes the behaviour of the two examined taphole clays.

Workability

and MEP of the taphole clay

Workability decay curves of taphole Clays A and B at 35°C, together with their workability specifications, are shown in Figure 2. Both clays exhibited an initial decrease in workability, with Clay B reaching a plateau at 33% (starting from 36%) after 14 days. In contrast, the workability of Clay A declined to 26% (also starting from 36%) after 21 days, corresponding to a 28% reduction. Clay B showed a workability loss of only 8%. These results suggest that the resole resin in Clay A significantly reduces the workability of the clay.

The MEP results (Figure 3) show a trend similar to that of the workability results. Over a period of 21 days, the load per unit area required to push the clay through an orifice increased for both Clay A and Clay B, with Clay B starting to plateau after 7 days. In contrast, the MEP of Clay A increased past 21 days. The constant increase in MEP of Clay A is indicative of a possible problematic operation of a mud gun as the taphole clay is pushed through with increasing difficulty. The samples for both workability and MEP were tested according to non-standard procedures. Based on experience of previous measurements the variability in the technique is approximately 2%−5% for workability and 5%−10% for MEP (Cameron, Garbers-Craig, 2025).

Binder viscosity changes

The workability decay results of clays with different binder viscosities are shown in Figure 4. These results indicate that, as the binder viscosity increases, the workability decay of the taphole clay increases. This confirms that lower-viscosity binders facilitate easier filling of the tap hole, although this does not account for the performance of the clay as a refractory material. A clear difference in workability reduction is observed between Clay A, which contains both resin and liquid pitch, and Clay B, which contains only liquid pitch (Figure 2). These results suggest that a mechanism contributing to the workability decay in resin-pitch mixtures likely involves an increase in viscosity.

Figure 2—Workability ageing of Clays A and B (R =Resole resin; P = Liquid pitch; Test temperature 35°C

Figure 3—MEP of Clays A and B (R = resole resin; P = liquid pitch; Test temperature 35°C)

Figure 4—Workability ageing of Clay A at different binder viscosities

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

Properties of the aggregate and matrix raw materials

Wettability of the solid particles by polar and non-polar liquids

The wettability results for samples tested in both polar (water) and non-polar (hexane) media are shown in Figure 5. The gradients from Figure 5 for both hexane and water, as well as the calculated angles, are presented in Table 2. All calculated contact angles are less than 90°, indicating that all raw materials, both aggregate and matrix material, are wettable by water and are therefore considered to have polar surfaces. This implies that the system behaves as a homogenous viscoelastic suspension of aggregate and matrix materials within a pitch (and resin) medium. Consequently, segregation of filler material can be eliminated as a potential cause of increased viscosity.

Chemical compositions and mineralogy of the aggregate and matrix materials

The chemical compositions and mineralogy of the aggregates and matrix raw materials were determined using XRF, XRD, and SEM-EDS analyses. The calcined alumina was highly pure (>99 mass% Al2O3), with only a trace amount of Na2O (0.14 mass%) as an impurity component. The andalusite had an alumina content of 62 mass%, with leucite (KAlSi2O6) as the main impurity. The bauxite contained 89 mass% Al2O3, with quartz and rutile (TiO2) as the main impurities. The calcined clay contained corundum and cristobalite as main phases, with muscovite ((KF)2(Al2O3)3(SiO2)6(H2O)) and rutile as impurity phases. The silicon carbide sample had a silicon carbide content greater than 90%.

The chemical composition and mineralogical analyses confirmed that none of the raw materials contained free lime or sulphur; components known to contribute to cross-linking of the binders used in taphole clay.

Characterisation of the liquid pitch, resole resin, and pitchresin mixtures

The results reported in this section confirmed that there is an interaction between the resin and liquid pitch that not only alters the structure of the resin but also changes the viscosity of the resinpitch mixture. These structural changes were investigated using FTIR, TGA, and DSC analyses.

Viscosity

The influence of shear rate on dynamic viscosity at 30°C was determined for the liquid pitch, resin, and resin-pitch combinations

Figure 6—Shear rate vs. viscosity rheology test of liquid pitch (P), virgin resole resin (R) and combinations thereof after mixing and ageing at 45°C. (Tests conducted at 30°C)

(Figure 6). The samples were tested after mixing and subsequent ageing at 45°C for two weeks. Shear rate tests revealed that both virgin resole resin and liquid pitch exhibit Newtonian behaviour. However, the age of resin-pitch mixtures displayed shear-thinning, non-Newtonian behaviour. The tendency of the zero-shear viscosity of the mixtures to exceed that of the virgin components clearly indicates an interaction between the resin and pitch. The extent of shear thinning increased with increasing resin content in the mixtures. This behaviour suggests a structural change wherein the polymer network of the resin breaks down into aggregate polymer chains or smaller polymer fragments (Gotro, 2018). The aged sample containing 75 mass% resole resin and 25 mass% liquid pitch could not be evaluated due to its excessively high viscosity. The viscosity of the virgin binders and resin-pitch combinations as a function of temperature and time (for both mixed and aged samples) was evaluated after mixing and after ageing at 45°C for two weeks (Figure 7). The temperature range used for evaluation was 30°C–130°C to capture the gel point of the virgin resole resin

2

Wettability constants and data used to calculate the contact angle between hexane (non-polar) / water (polar) and different aggregate and matrix materials

Figure 5—Wettability results for aggregate and matrix raw materials used in Clays A and B (P = polar medium, NP = non-polar medium)

Table

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

in viscosity with an increase in temperature of resole resin, liquid pitch, and combinations thereof (shear rate = 1 sec-1)

at approximately 125°C. The viscosity of the virgin liquid pitch decreased with an increased temperature up to 110°C, after which it began to plateau. In contrast, the viscosity of the virgin resin decreased up to 118°C, followed by a sharp increase between 12°C and 126°C, consistent with the known curing region of resole resin (Wilmer, 2014; Gotro, 2014). Upon ageing, the samples exhibited an increase in initial viscosity, confirming that ageing contributes significantly to increased viscosity and, consequently, to reduced workability.

In the mixed (not aged) resin-pitch samples, no distinct increase in viscosity was observed within the 30°C–130°C range, unlike the virgin resole resin. This suggests that the presence of liquid pitch interferes with the resin structure required to initiate cross-linking, a typical behaviour in thermosetting resins (Fan, Weclawski, 2017). The viscosity of aged samples increased significantly, more so than that of the mixed samples, particularly in the 70°C–130°C range. As the resin content increased, the magnitude of the viscosity increase became more pronounced, presumably due to cross-linking starting at lower temperatures as a result of the higher reactive resin fraction.

The aged sample containing 25 mass% resole resin and 75 mass% liquid pitch exhibited multiple viscosity increases, one at 75°C–95°C and another between 115°C–130°C, which can be attributed to multiple curing reactions (Strzelec et al., 2012). This indicates that the onset of curing occurred at lower temperatures. Increasing the resin content also resulted in multiple curing reactions, some of which likely shifted to temperatures above 130°C.

Thermogravimetric analysis (TGA)

The rheological observations were supported by TGA, DSC, and FTIR analyses. TGA was used to evaluate the mass loss of the virgin resin and virgin liquid pitch samples in the temperature range of 30°C–130°C (Figure 8). The mass of both the resin and the pitch remained constant up to approximately 50°C. Beyond this temperature, the pitch began to exhibit mass loss, while the mass of the resin remained virtually unchanged up to 100°C. Above 100°C, a slight reduction in resin mass was observed, possibly due to the onset of water release associated with the gelation process. However, this minor mass loss is unlikely to significantly affect the viscosity changes observed in the pitch-resin mixtures used in the taphole clay.

Differential scanning calorimetry

Differential scanning calorimetry (DSC) analyses were conducted from 20°C to 130°C in air, to simulate the reactions that occur during manufacturing (Figure 9). The peak at approximately 106°C for the virgin resole resin, corresponds to a mass loss at ~100°C observed in the TGA curve. The curing reaction occurred between

65°C–120°C and was marked by an exothermic peak (Haddadi et al., 2017).

The DSC curves of the virgin liquid pitch and the mixed resinpitch combinations did not show any peaks during heating. This observation aligns with the viscosity results for the resin–pitch mixed samples, where no abrupt changes in viscosity were observed as the temperature increased. Both the DSC and viscosity data confirm that no curing reactions occurred in the resole resin-liquid pitch mixtures that were only mixed. This suggests that the structure of the resin is altered upon mixing with the liquid pitch, thereby preventing typical curing behaviour.

The aged resin-pitch mixtures however, exhibit distinctive exothermic peaks at various temperatures. As the proportion of resole resin in the mixtures increased, the exothermic peaks shifted to lower temperatures. By comparing these DSC results with the viscosity data of the aged resin-pitch samples, it is evident that curing of the resin-pitch mixtures occurs at lower temperatures than in the virgin resole resin. This is presumably due to a cross-linking interaction between the resole resin and the liquid pitch after mixing, which alters the structure of the resole resin. During ageing, the resole resin begins to cross-link prematurely.

Fourier-transform infrared spectroscopy – ageing and curing test

Fourier-transform infrared spectroscopy (FTIR) spectra of the liquid pitch, resole resin as well as the 1% resin – 17% pitch combination used in Clay A are shown in Figure 10. The spectrum of Clay A represents the unaged 1% resin – 17% pitch mixed. The identified functional groups are reported in Table 3. The structure of the resole resin was confirmed to be that of a phenol-formaldehyde structure (Authier-Matrin et al., 2001). Key features include a broad O-H peak at 3250 cm-1, an SP2 C-H stretch at 3050 cm-1, a cyclic aromatic structure peak at 1610 cm-1, and C-O-C bonds between 1000 cm-1200 cm-1

Figure 7—Changes

Figure 8—TGA analysis of virgin resole resin and liquid pitch binder (nitrogen atmosphere)

Figure 9—DSC results for liquid pitch, virgin resin, and resin-pitch combinations (mixed and aged in air)

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

Table 3

Interpretation of FTIR results for liquid pitch binder, resole resin, and resin-pitch combination used in Clay A

Observed wavelength (cm-1)

~3070

~2900

Wavelengths from literature (cm-1)

3000–3300

Functional groups

2850–3000 N-H bond (R2NH)

~1605 1600 SP3 C-H stretch

~1180 1180 C=C stretch (aromatic)

~750 (fingerprint)

C-N stretch

~730 (fingerprint) 737 Benzene C-H bending

Resole resin N-H wag (R2)

~3250

~3050

~2920

3200–3600

3000–3150

O-H bonds (broad peak)

2850–3000 SP2 C-H stretch (hybridised) – peak overlaps with O-H

~1590 1600–1650 SP3 C-H stretch

~1200 (fingerprint)

~1033 (fingerprint)

1000–1200 (multiple peaks) C=C stretch (aromatic)

1000–1200 C-O-C bonds

Figure 10—FTIR identification of functional groups associated with the liquid pitch, the resin, and resin-pitch mixture (1% resin – 17% pitch) used in Clay A (before ageing)

The FTIR spectrum of the liquid pitch in Figure 10 shows the presence of aromatics and SP C-H stretches. Due to the complex structure of the liquid pitch, overlapping peaks from various compounds are present. Distinctive functional groups such as secondary amines are identified by the N-H stretch (at ~3070cm-1), C-N stretch (at ~1183cm-1) and N-H wag (at 737cm-1).

The spectrum of the resin-pitch mixture used in Clay A does not exhibit any major peak changes. It closely resembles that of the liquid pitch, with a minor change in the O-H bond region, which can be contributed to the presence of the resole resin in the sample.

To further investigate the effects of mixing, ageing and curing of the resole resin and liquid pitch, changes in viscosity and functional groups were evaluated using FTIR/ATR and DSC.

The possibility of early-onset cross-linking was evaluated through curing tests on the individual binders and corresponding FTIR analyses. The curing test involved heating the samples at 130°C for 2 hours to ensure complete curing. The result for the virgin resole resin is shown in Figure 11. During curing, free O-H bonds appeared at ~3680 cm-1 (Figure 11(a)), the SP3 C-H stretch at ~2920 cm-1 increased (Figure 11(b)), while the C-O-C stretch at ~1020 cm-1 increased significantly (Figure 11(c)), all of which indicate a higher cross-linking density during the curing process (Zhou et al., 2010).

12—FTIR curve of virgin liquid pitch and liquid pitch

(arrows indicating the amine peaks)

The curing test result for the virgin liquid pitch is shown in Figure 12. As expected, no changes in functional groups were observed during curing, confirming the inertness of the virgin liquid pitch under the test conditions.

A second set of FTIR tests was conducted to identify functional group changes in samples that were mixed, aged, and cured. The mixed and aged test results correspond to the same samples used for viscosity and DSC analysis. The ageing and curing test results

Figure 11—FTIR curve of virgin resole resin and resole resin cured at 130°C (a = free O-H bonds at ~3680 cm-1, b = SP3 C-H stretch at ~2920 cm-1, c = C-O-C stretch at ~1020 cm-1)

Figure

cured at 130°C

Premature ageing of a blast furnace taphole clay containing resole resin and liquid pitch

for the 50 mass% resin – 50 mass% pitch mixture is given in Figure 13. A significant increase in the peak intensity of the C-O-C stretch functional groups at ~1020 cm-1 was observed. As the resin content increased, these C-O-C stretch functional group peaks in the aged samples also became more prominent. The increases in C-O-C functional group intensity between the mixed and aged samples is attributed to the cross-linking of the resole resin in the presence of liquid pitch (Anis et al., 2011; Dole, 1979).

There were no shifts in peak positions, only changes in the intensity of the C-O-C peaks. Peaks associated with secondary amines were present in the mixed, aged, and cured samples. In all three samples, the N–H peak at ~3070 cm-1 remained visible in the mixed and aged samples. In the cured sample, this peak was not easily identifiable due to overlap with a strong peak at 3000 cm-1 However, its presence cannot be ruled out. The C-N stretch and N-H wag peaks, located at ~1180 cm-1 and ~737 cm-1, respectively, were present in all three samples.

It is likely that the presence of secondary amines in the liquid pitch acts as a catalyst for the cross-linking of the resole resin during ageing, thereby accelerating the curing process. The FTIR results of the aged and cured samples confirm that, as the mixtures aged, the resin began to cross-link, resulting in an increase in the peak intensity of the C-O-C stretch functional group. This increase is most likely due to acetal ring formation, which is attributed to premature cross-linking of the resole resin catalysed by the secondary amines in the liquid pitch (Anis et al., 2011).

Conclusions

Causes for reduced workability and increased ageing of a blast furnace taphole clay were examined in this study. The following conclusions can be drawn:

➤ The workability of the investigated taphole clay decreased with time when 1% resole resin and 17% liquid pitch were used as binders, compared to using 18% liquid pitch alone. The reduction in workability was associated with an increase in binder viscosity.

➤ The MEP of the investigated taphole clay increased over time when resole resin and liquid pitch were used in combination, as opposed to using only liquid pitch.

➤ The surface chemistry of the aggregate and matrix raw materials was polar and compatible with that of the resole resin and liquid pitch used as binders. No impurities, such as sulphur or free lime, were detected in the aggregate or matrix raw materials that could have contributed to premature crosslinking of the resin.

➤ Mixtures of resin and pitch exhibited higher viscosities than either pure resin or pure liquid pitch and exhibited shearthinning behaviour.

➤ The viscosities of resin-pitch mixtures increased with ageing due to premature cross-linking of the resin structure. C-O-C bonds between polymer chains became more prominent over time and with higher resin content. Functional groups of secondary amines were identified in the molecular structure of the liquid pitch.

➤ Premature cross-linking in the binder mixtures is likely caused by the secondary amines in the liquid pitch, which act as catalysts for the curing reaction of the resole resin. This, in turn, resulted in a lowering of the curing temperature of the resin-pitch mixtures.

➤ The observed reduction in workability of the taphole clay was therefore attributed to increased binder viscosity during ageing, caused by premature cross-linking in the binder system and the reduction in curing temperatures when resole resin was mixed with liquid pitch.

Acknowledgements

Funding from the Anglo American Chair in Pyrometallurgy, Department of Materials Science and Metallurgical Engineering at the University of Pretoria, is gratefully acknowledged.

Author credit statement

IJ-PC: Conceptualisation, investigation, writing original draft; SR: Supervision, writing - reviewing and editing of draft; AGC: Supervision, funding acquisition, Writing - reviewing and editing.

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SAIMM URANIUM CONFERENCE 2026

INTRODUCTION

0ne kilogram of uranium can produce as much energy as 160 tons of coal. As the world transitions to sustainable, low-carbon energy solutions, uranium will play an increasingly critical role by enabling the generation of large amounts of electricity with minimal greenhouse gas emissions. Uranium already forms part of a reliable and low-emission energy mix in many countries, contributing significantly to global decarbonisation efforts.

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FOR FURTHER INFORMATION, CONTACT:

Gugu

Charlie, Conferences and Events Coordinator

E-mail: gugu@saimm.co.za

Tel: +27 11 530 0238

Web: www.saimm.co.za

CONFERENCE PROGRAMME

Papers are invited on the following topics:

• Uranium market trends

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• Mining

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• Refining and value-added products

• Fuel cycle

• Recycling and reprocessing

• Nuclear/radioactive waste and site remediation

• Logistics of handling and transporting uranium in its various forms

• Medical applications

• Health and safety

• Environment, Social and Governance (ESG)

• Legislative and policy issues

• Economics

KEY DATES

• 2 March 2026 - Submission of abstracts

• 13 April 2026 - Submission of papers

• 18 August 2026 – Technical workshop: Modelling with Cycad Process

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Harmony Gold mining company – South Uranium plant, ion-exchange plant

Langer Heinrich Uranium

Affiliation:

1Department of Mining Engineering, University of Pretoria, South Africa

Correspondence to:

I. Maubane

P.L. Ngwenyama

Email: ingridseipati14@gmail.com larrance.ngwenyama@up.ac.za

Dates:

Received: 30 Nov. 2023

Revised: 25 Jul. 2025

Accepted: 15 May 2025

Published: July 2025

How to cite:

Maubane, I., Ngwenyama, P.L. 2025. The use of soundless chemical demolition agents in large scale in situ rock breaking applications in the mining industry. Journal of the Southern African Institute of Mining and Metallurgy, vol. 125, no. 7, pp. 371–384

DOI ID:

https://doi.org/10.17159/2411-9717/3204/2025

ORCiD:

I. Maubane

http://orcid.org/0009-0003-1275-4352

P.L. Ngwenyama

http://orcid.org/0000-0002-9568-4964

The use of soundless chemical demolition agents in large scale in situ rock breaking applications in the mining industry

by I. Maubane1, P.L. Ngwenyama1

Abstract

Soundless chemical demolition agents (SCDAs) are proving to be the future of sustainable and environmental-friendly mining. These expansive chemicals become critical when there is a need to break rock close to communities, environmentally sensitive areas, and critical structures and infrastructure. Unlike explosives, the SCDAs do not generate air-blast, ground vibrations, flyrock, noise, dust, and noxious fumes, which can negatively affect the environment, surrounding communities, sensitive areas, and critical structures and infrastructure. Over the years, several types of SCDA products have been used successfully in the civil and construction industries for the demolition of old and dilapidated buildings near vibrations and environmentally sensitive areas. They have also been used to fragment boulders in the mining and construction industries. However, to date, no study has investigated the applications of soundless chemical demolition on a large-scale in opencast mines. Therefore, this study evaluated whether soundless chemical demolition can break large volumes of in situ rock. This was done by conducting five trials using Nex-Pand soundless chemical demolition agents in four different surface mining sites. Trial 1 was unsuccessful as it was unable to break the rock due to the large confinement between the drilled holes and the use of a hole diameter that was 182% larger than recommended. Trial 2 was a success as it was able to fracture the rock with a crack width that reached 110 mm after 72 hours. Trial 3 generated cracks at the biggest hole diameter of 102 mm. The cracks developed and ceased to increase after 96 hours at a crack width of 100 mm. Trial 4 achieved the smallest crack width due to a very competent rock compared to the other sites. Trial 5 generated cracks at a slower rate than other trials because of the use of Nex-Pand powder at lower ambient temperatures than recommended. These trials proved that soundless chemical demolition agents can replace explosives in areas closer to sensitive infrastructure and communities. This will ultimately enable mines to unlock value and access deposits in these areas.

Keywords

non-explosive rock breaking, soundless chemical demolition, crack development, crack growth and fragmentation

Introduction

For centuries, explosive rock breaking has been an integral part of the mining value chain and has been extensively used across all types of mines (Dzimunya et al., 2023). The use of explosives has been one of the most efficient and cheapest methods for breaking in situ rock in order to access minerals of economic value. Despite all the successes, explosive rock breaking has had some inevitable challenges and disadvantages. Only 20% – 30% of the energy released during detonation is converted into mechanical energy that actually fragments the rock (Malbašić, Stojanović, 2018). The remaining 70% – 80% of the explosive energy is dissipated in the form of excessive noise, ground vibrations, back breaks, air-blast, and fly-rock (Malbašić, Stojanović, 2018). These factors have a negative impact on the health and safety of personnel, the surrounding communities, and the environment, and they can cause damage to property, equipment, and infrastructure (Zhou et al., 2018). For example, communities may experience effects such as broken windows and cracked walls due to excessive ground vibrations and air-blast. Due to these effects, legal limitations have been enacted to regulate the use of explosives in areas closer to property, structures and infrastructure, and surrounding communities. Regulation 4.16 of the Mine Health and Safety Act (MHSA No. 29 of 1996) stipulates that there should not be any blasting operations within a horizontal distance of 500 m from residential structures unless permission has been obtained from the principal inspector of mines. Before the permission can be obtained, the mine is required to conduct a risk assessment and consult with the stakeholders that would potentially be affected by the blasting activities. These limitations are currently imposed on one of the coal mines

The use of soundless chemical demolition agents in large scale in situ rock breaking

operating in the Witbank Coalfield (Mine A). Mining activities at this mine are quickly approaching coal reserves situated closer to a community, a farm, and a national road. As a result, this has introduced a new challenge for the mine as the use of explosives for blasting was restricted. This can be seen in Figure 1, where one of the mining areas, referred to as Block 10, has advanced very close to a community, a farm, and a national road. Consequently, coal reserves from the Block 10 area are currently unmined due to blasting zone restrictions imposed by the MHSA.

The Block 10 area is located approximately 66 m from the first house in the nearby community. Figure 2 is a depiction of the 500 m boundary (red circle) between Block 10, the surrounding community (Re 4 and Re 17), and the national road. The solid purple block shows the Block 10 (Re 10) area, which is planned and scheduled for production. The green lines show the mine boundary, the brown lines show the road infrastructure, and the blue lines show a river stream.

It can be seen in Figure 2 that a significant number of community houses and the national road are within the 500 m boundary, hence permission is required before Block 10 can be blasted. The permission to use drilling and blasting to break the rock at Block 10 was not granted at the time of the study, thus the coal resources in Block 10 are still unmined. Block 10 is comprised of three coal seams, namely, the No. 3 seam, the No. 2 seam, and the No. 1 seam. The geology of Block 10 consists of coal, sandstone, and mudstone on the overburden and interburden. The No. 2 coal seam is 5.87 m thick, and 14.54 m deep. It is currently the only viable seam planned to be mined at the current conditions. The uniaxial compressive strength (UCS) of the area ranges from 30 MPa to 74.9 MPa. Coal has the lowest UCS whereas overburden has a high UCS. The No.2 seam has 2.03 million tonnes. These coal resources cannot be accessed using the current rock breaking method. Although the study was based at the coal mine mentioned (Mine A), various other mining sites facing similar challenges formed part of the study in which further experiments and tests were conducted. These sites were unable to conduct normal blasting due to being closer to sensitive infrastructure. This includes a coal mine (Mine B), which could not blast due to the coal deposit being situated within 500 m to a community. Mine C, a platinum mine could not remove in situ rock for the development of slurry dams as the rock is within 50 m of plant infrastructure. Similarly, another platinum mine (Mine D),

could not remove in situ rock for the purpose of developing a slurry dam due to being closer than prescribed to a processing plant. These challenges prompted an industry-wide need to find and evaluate non-explosive rock breaking methods that can be used at a large scale when blasting closer to sensitive structures and infrastructure. As such, several types of non-explosive rock breaking methods or techniques were considered and evaluated for their applicability on a larger scale mining operation. These methods include hydraulic impact rock breaking, thermal rock breaking, soundless chemical demolition agents (SCDAs), controlled foam injection, and mechanical splitters (rock splitters). These methods were evaluated based on the safety of the method, the environmental effects, and applicability to the geological conditions at the mining sites.

During hydraulic rock breaking, high-pressure steady or pulsejet water is injected into drilled holes. The high-pressure water supplies the energy required for shattering and cutting rock. This method does not generate any harmful or toxic gases, is non-explosive or destructive, and thus, safer than explosives. Furthermore, it is more effective in splitting unconfined boulders, rather than breaking rock, under confined conditions (Singh, 1998; Genet et al., 2009). The controlled foam injection method makes use of a high-pressure foam (gas and liquid), which is quickly poured to the bottom of the hole to create expansive pressure (Young, Graham, 1999). Although the foam is inert and expands quickly, its use may lead to the generation of fly-rock and air-blast (Singh, 1998). This method is mostly used on short and smaller diameter holes (De Graaf, Spiteri, 2018). The thermal spalling rock breaking method uses high temperatures to weaken and shatter the rock (Yang et al., 2024). The heat is required to erode the surface layer of the rock in the holes. This method is also environmentally friendly; however, it cannot be used to fragment rocks such as sandstone, as an endothermic reaction occurs during heating. The mechanical splitter rock breaking method makes use of a wedgeset that is inserted into drilled holes to generate fractures, which ultimately breaks the in-tension (Li et al., 2024). This method does not generate noise, dust, vibrations, and shock. It is most effective with the presence of a free face and on a typical hole diameter of 63.5 mm and hole depth of 381 mm. It is relatively expensive and generally used in large scale underground applications with limited depth (De Graaf, Spiteri, 2018; De Graaf, 2018). The SCDA method makes use of chemicals that are mixed with water, poured into

Figure 1—Block 10 location relative to houses from the nearby community

Figure 2—Block 10 location at 500 m boundary to the nearby community

The use of soundless chemical demolition agents in large scale in situ rock breaking

drilled holes, and allowed to expand over time (Zhong et al., 2023). The SCDA rock breaking method happens during a slower process, making it safer to use and handle. This improves safety by reducing the risk of accidents and injuries. The SCDAs produce minimal noise, vibrations, and dust. They do not produce harmful fumes and fly-rock (Al-Bakri, Hefni, 2021). Furthermore, the chemical product allows for a controlled rock demolition or fracturing process, which minimises the potential for damage to surrounding structures (Natanzi et al., 2016). The chemical product can be applied in different geological mining and construction sites due to their versatility and can be used to break different rock types. They do not require specialised labour, safety measures, and cleaning up after a blast (Maneenoi, Bissen, Chawchai, 2022).

After the evaluations and comparisons of the different nonexplosive rock breaking methods, the use of SCDAs was selected as a potential and most suitable method to use for the trial due to their safety, minimal environmental effects, and their versatility to fragment different rock types and be applied in different geological mining and construction sites. However, the SCDAs have not yet been tested to break large volumes of in situ rock. To the best of our knowledge, no study has published research on the application of the SCDAs on a large-scale open-cast mining operation, both locally and globally. Some work has been conducted in underground mines (Habib, Shnorhokian, Mitri, 2022). Therefore, the purpose of the study was to evaluate the applicability of the SCDAs to fragment large volumes of in situ rock. This was done by conducting five trials of SCDAs on four different surface mining sites using Nex-Pand powder. The choice of SCDAs as opposed to the other non-explosive rock breaking methods for this study was motivated by their safety, minimal environmental effects, and applicability to the geological conditions at the mining sites. This proved to be a viable solution for fragmenting rock in sensitive structures and infrastructure. However, the SCDAs have not yet been tested to break large volumes of in situ rock.

Literature review

The SCDAs, often referred to as demolition agents or expansive cements, are powdery substances that expand when mixed with water. The expansion occurs through a chemical hydration process, by the formation and development of ettringite crystals. The SCDAs are mainly used in the mining and construction industries to break concrete or boulders of other types of rock. However, the SCDAs failed to gain mainstream adoption for the selective removal of rock and concrete when they were first introduced in the early 1970s (Al-Bakri, Hefni, 2021). This was due to their proprietary nature (privately owned and controlled) and a lack of guidelines on their usage. Currently, the patents have expired, and more competitive products have entered the market.

Administration process

SCDAs make use of drilled holes. Holes of a certain diameter and depth are drilled in the material or rock to be broken. The SCDA is then mixed with water at a certain ratio depending on the manufacturer's specifications. The slurry is poured into the drilled holes and left to solidify. The solidification period or setting time of the slurry varies according to the type of demolition agent. During the solidification process, the slurry expands in volume. The expansion of the slurry pushes against the walls of the hole and thereby causes cracks to develop on the walls (Gómez, Mura, 1984; Xu et al., 2023). The expansion pressure or force is limited by factors such as the intensity and rate of the chemical reaction, the

rock confinement, chemical composition, accuracy of the chemical mix with water content, and the environmental or atmospheric conditions (temperature sensitivity) (De Silva et al., 2016). The maximum expansion pressure depends on the volume of the gas produced and the heat generated from the hydration of calcium oxide (CaO) to calcium hydroxide (Ca(OH)2) in the exothermic reaction. The hydration process can be affected by too little water, but excessive water dilutes the composition, thereby limiting the expansion pressure. The expansion pressure can also be limited when the temperature is too high or too low. The expansion pressure is necessary for generating the cracks between adjacent holes. The interaction mechanism of two neighbouring holes filled with SCDA is shown in Figure 3.

The expansion of the mixture generates pressure that acts on the inside wall of the drilled holes. The pressure results in compression of the rock between the holes, which causes tensile stresses perpendicular to the line connecting the two holes. The rock between the holes will develop fractures when the tensile stress produced during expansion exceeds the tensile strength of the rock being broken (Shang et al., 2018).

Chemical composition

The expansive cements are made of Portland cement and an expansive additive. There are four types of SCDAs. These vary based on the expansive additive. There are type K, type M, type S, and class G cements (Habib, Shnorhokian, Mitri, 2022). Type K, M, and S commonly produce ettringite crystals, which is one of the forces that drives the expansive capability. However, they each vary in the source of the aluminate component and the operating temperature (Habib, 2022). For example, type K operates optimally between -5°C and 10°C, type M at 10°C to 20°C, type S at 20°C to 35°C, and type G at 35°C and higher. This makes Class G ideal for the study and will be discussed as the SCDA powder used in this study. Class G is a type of expansive cement in which Portlandite drives the expansive capability (Arshadnejad, Goshtasbi, Aghazadeh, 2011).

Class G

This expansive cement constitutes 80% – 90% of CaO (lime) (Habib, 2022). Other elements such as silicon (SiO2), aluminium oxide (Al2O3), ferrous oxide (Fe2O3), calcium fluoride (CaF2), and magnesium oxide (MgO) have been added to improve, alter, delay, or control the hydration process (Natanzi et al, 2016). The addition of water to the demolition agent results in a hydration reaction with calcium oxide (CaO). The hydration generates heat and calcium hydroxide under the exothermic reaction (Natanzi et al, 2016). The chemical reaction is given by Equation 1 (Arshadnejad, Goshtasbi, Aghazadeh, 2011).

CaO + H2O → Ca(OH)2 + 15.2 kcal/mol [1]

Figure 3—Interaction mechanism of two adjacent holes subjected to the expansive pressure from SCDA (Shang et al., 2018)

The use of soundless chemical demolition agents in large scale in situ rock breaking

The hydration reaction of calcium oxide is the root of class G cement’s expansive force (Habib, 2022). A higher calcium oxide content causes more reaction and higher expansive pressures. High CaO content is the main cause of volumetric expansion. The hydration of lime can create up to 1.9 times the original volume. The generated expansive pressure is dependent on the degree of hydration of CaO. From the chemical reaction in Equation 1, the exothermic reaction creates 15.2 kcal/mol of energy. As a result of the exothermic property, CaO based SCDA can reach a heat of hydration of up to 150°C, such that the boiling water for curing/ solidification can potentially create blowouts (Soeda, Harada, 1994). The likelihood of blowouts occurring during field applications is unlikely as the heat is dissipated into the surrounding environment (Natanzi et.al, 2016). The generated expansion pressure of CaO or lime reaches 30 – 44 MPa. The pressure necessary to fragment soft rock or concrete is 10 – 20 MPa (Al-Bakri, Hefni, 2021).

Parameters affecting the effectiveness of the rock breaking process

Ambient temperature

SCDAs are designed to be applied over a vast range of ambient temperatures. The selection of the correct SCDA type depends on the lowest temperature possible in the area (Huynh, Laefer, 2009). Natanzi et al (2016) proved that high ambient temperatures reach the time to peak SCDA temperature development (hydration heat) faster and results in larger expansive pressures, more rapid pressure gain, and bigger volumetric expansion. The results showed that SCDAs are more successful when applied in the temperature range that they were designed under. Laefer et al. (2010) proved that higher ambient temperatures result in faster cracking and that SCDAs designed to work in cold environments can be used in areas of high ambient temperature to speed up the cracking process. The time to create the first crack varies from a few to almost 24 hours, depending on the hole parameters and the rock properties.

Water used for mixing

Usually, the water temperature should be less than 15°C with a ratio of water to SCDA being 1:3 by weight. The use of too little water will result in inadequate availability of water to correctly hydrate all the SCDA powder, while the use of surplus water will cause free water to remain and possibly compromise SCDA (Hinze, Brown, 1994). Thus, it is important to follow the manufacturer’s recommendations when mixing the SCDA with water. Laefer et al. (2010) proved that increasing the water temperature by 152% results in an 18.92% reduction of time to the first crack.

Hole parameters and rock strength

Laefer et al. (2010) proved that a larger burden and distance to free face result in an increased time to first crack and that specimens with artificial seals lead to a reduced time to first crack. A bigger hole diameter, within the manufacturer’s recommendations, generates more pressure (Hanif, 1997). Shang et al. (2018) proved that, for the same hole diameter, increasing the hole spacing increases the time that it takes to fracture the rock. Given the same parameters, stronger materials require a closer hole spacing to get similar cracking levels (Arshadnejad et al., 2011). Chemical admixtures improve the rate of stiffening, setting, hardening rate, or early strength (Fu et al.,1995).

Current application of SCDAs in the mining industry and related fields

To date, there is very limited research having been conducted on

the applications of the SCDAs in the mining industry. This is the first paper to present experiments or trials of actual findings. But then there is also limited work in related industries that presents field trials of the SCDAs. The SCDAs were initially developed for the construction and civil engineering industries for rock breaking applications where explosives could not be used due to safety, environmental, and regulatory constraints. This included breaking rock close to sensitive infrastructure and demolition of old and dilapidated buildings. Some of the notable work has been focused on reviewing the SCDAs and conducting laboratory experiments to enhance their performance in the construction industry (Hinze, Brown, 1994; Hinze, Nelson, 1996; Gambatese, 2003; Huynh, Laefer, 2009; Habib, 2019; Kim et al., 2021; Zhong et al., 2023). The product has been a success in the demolition of old and dilapidated structures such as hospitals, monuments, schools, buildings near residential buildings, cities, and urban areas. These areas are highly sensitive to vibrations, noises, toxic fumes, dust, and fly-rock where the use of explosives is highly restricted. Some of the SCDA experiments and applications have been focused on breaking reinforced concrete structures without any damages or minimal damages (Jiang et al., 2021; Jiang et al., 2022; Li et al., 2023; Jiang et al., 2023). In mining, limited research work has been conducted and has focused on conducting laboratory experiments for future underground mining applications (Habib, 2019; Habib, 2022; Habib, Vennes, Mitri, 2022; Zhong et al., 2023). Some tests have been conducted for granite and sandstone blocks, which can be associated with fragmenting boulders in opencast mining operations (Sakhno, Sakhno, 2024; Yapici Tanyeri, 2023). These studies have been limited to experiments in laboratories without real life applications. As such, this study aims to conduct actual experiments or tests on actual mining blocks on a larger scale rather than the laboratory experiments.

Methodology

Before this study, SCDAs were only used in the construction industry, near structure demolition applications, small-scale laboratory experiments, and underground tunnelling rock breaking applications. However, to date, no study has tested or experimented on the application of the SCDAs in large-scale rock breaking applications. Therefore, this study was required to first formulate a methodology for the process of fragmenting rock using the SCDAs on a large-scale application. Therefore, the first experiment or trial was used to document the methodology to be used in the other trials. The methodology was developed in the steps illustrated in Figure 4 for opencast mining.

After the block had been prepared, the required holes were drilled using a Pantera 1500id drill rig. The process of pouring the chemical into the drilled holes commenced. Figure 5 shows the steps that were followed to mix and pour the chemical into the holes. The aim was to establish a relationship between the crack development and the amount of powder. The development of the cracks can be measured by its length and width development. This will determine the capability of the SCDA to fragment the rock.

Results and discussion of results

In this study, five trials were conducted using two types of Nex-Pand powder. All the powders were of Extreme Summer type. The first two trials made use of a powder designed to work on 30 – 50 mm hole diameters and the last three on 89 – 130 mm hole diameters. More information on the small hole chemical can be found on the company’s website (https://www.harlensupplies.co.za/). Information about the large hole chemical is not yet available to the public.

The use of soundless chemical demolition agents in large scale in situ rock breaking

Firstly, the site was selected. When selecting the area, it was necessary to ensure that the area was dry and had no ground water. The requirement for the area that the SCDA is going to be used, is that the area should have no/ less groundwater. Water should not fill more than 15% of the hole to prevent possible alterrations to the slurry mixture (D. Harlen, Personal Communication, 10 January 2023). The ambient or hole temperature of the area should match the temperature that the selected SCDA was designed to operate in. In this experiment, the temperature of the area should be between 20ºC and 40ºC.

The area was then prepared for drilling using a tyre dozer to clean, level the area and create access roads and ramps into the block.

After cleaning, the presence of cracks was assessed. Cracks could cause runoff of the chemical inside the hole or cause the holes to close, thus affecting the effectiveness of the chemical. Polythylene sheath was suggested to prevent the run-off of the product due to cracks. The sheath was to be placed inside the hole before pouring the chemical in an area with cracks to prevent run-off. However, the polythylene sheath was not available at the mine at the time of the study. For control measures, the depth remaining after pouring the chemical into the holes was measured and compared for all holes. This was to check if there were any run-offs.

the hole depth: This quantity of powder to be poured into each hole depends on the hole depth. The depth was measured to ensure that the correct amount of Nex-Pand power chemical was then poured into each hole. The quantities poured into the holes were based on design factors such as the gold diameter, depth of holes, burden and spacing, rock strenghth, etc. These quantities differed for the different trials at the different mining sites.

4. Mixing the product: The Nex-Pand powder was then mixed with the required amount of water as per the specific site conditions such as the hole diameter and depth. The product was interminably mixed until it was poured into the holes using the buckets. The holes were not stemmed (were left uncovered).

Figure 5—Preparation for filling the drilled holes with chemical

Results of Trial 1 conducted at Mine A

This trial was conducted on sandstone type rock. Table 1 shows the parameters used in this trial.

All the holes had a remaining depth of 1 m after pouring. This proved that there were no run-offs after pouring. If there were any run-offs, the resulting depth of the holes would have varied. The small hole Nex-Pand powder used was designed to work on diameters ranging between 30 and 50 mm. The experiment started at 17:00. The holes were left overnight and checked after 15 hours, which was at 08:00 of the following day. The trial was only conducted on four holes to evaluate if this chemical has the potential to break the rock and to document the methodology.

Results after 15 hours

Cracks indicating fracturing were expected. The cracks were to be evaluated in terms of width and length. The crack length from each

hole was expected to grow and reach that of the adjacent hole. The growth in length of the cracks was to be in correlation with the growing width of the cracks. The chemical blew out of the holes without generating any cracks.

Results after 48 hours

Although the product in the holes started blowing out, this was a sign of the expansiveness of the product inside the holes. Therefore, it was decided to allow more time to see if the product remaining in the holes would yield any positive results. The holes were then checked at regular time intervals. However, after 48 hours it was realised that there were still no visible cracks between the holes and at the collar of the hole. This suggested that the product had the potential to break the rock but was too confined due to the burden and spacing used. There was not enough compressive stress between the rocks to generate tensile stress that exceeds the tensile strength

Figure 4—Site preparations process at site (steps followed before drilling

1. Measuring the temperature of the water: This was done to ensure that the temperature matches the recommended water temperature of less thank 15°C.

2. Measuring the hole temperature: This was done to ensure that the holes meet the temperature requirements of the Nex=Pand powder used in the trials.

3. Measuring

The use of soundless chemical demolition agents in large scale in situ rock breaking

Table 1

Trial 1 parameters Parameter

Distance from the edge/free face

Hole spacing 1 m

Number of holes poured with the agent 4

Block size of poured holes 1 m2

Quantity of powder and water mixed in hole 40 kg of powder (8 bags) and 12 L of cold water (15°C)

Time taken to mix and pour 4 holes 30 minutes

Labour Two blasting assistants

Depth remaining after pouring 1 m

Table 2

Factors affecting crack development in Trial 1 Parameter

of the rock. There was a greater rock to compress between the holes. The results from this trial are shown in Figure 6. There were no cracks generated during this trial and thus no fracturing took place. The trial did not yield the desirable results. Factors explained in literature were evaluated to determine the parameters that led to the generation of no cracks in this trial. The factors are summarised in Table 2.

It was found that the actual hole diameter used in the trial was 182% bigger than the theoretical diameter. There were no additions of chemical admixtures to the chemical during the trial. The parameters used correctly did not lead to undesirable results. The

Produced expanding pressures between 40 and 100 MPa

trial proved that this chemical has the potential to break the rock, however, a bigger hole diameter and great confinement between the holes led to no crack generation.

Trial 2: Mine B

This trial was conducted on dolomite rock with a tensile strength ranging between 6.2 MPa and 27.4 MPa. This trial was conducted on reduced confinement and hole diameter. The block size was increased to evaluate if the chemical would be able to fragment large volumes of in situ rock. Table 3 illustrates the input parameters during the trial.

Figure 6—Trial 1 results from site

The use of soundless chemical demolition agents in large scale in situ rock breaking

The hole diameter drilled in this trial is 28% bigger than the maximum recommended 50 mm. The hole depth, chemical type, and powder quantity was kept constant as in Trial 1. The chemical fractured the rock. The first crack became visible after 3 hours (with a crack width of 5 mm), shown in Figure 7. The crack width stopped increasing after 72 hours when it reached a crack width of 110 mm.

The crack length reached saturation and stopped increasing after 24 hours, as shown in Figure 8. The crack length reaches saturation when it extends from one hole to the adjacent hole.

Figure 9(a) shows the type of fracturing observed and the cracks generated by the expansion after 24 hours. The outcomes of the fracturing were measured through the resulting fragmentation. This was to ensure that the broken material could be loaded efficiently. Figure 9(b) shows the resulting fragmentation after the fracturing process was completed. The fragmented rock was loaded by an excavator with a 2.2 m3 bucket capacity a week later. The period that went by before loading was due to the availability of the machine.

The chemical fragmented the rock at a reduced confinement and hole diameter (28% bigger than the manufacturer’s specifications) compared to Trial 1. The water quantity used for mixing, mixing time, ambient temperature, and type of Nex-Pand powder were used according to specifications.

Trial 3: Mine A

After the unsuccessful trial at Mine A, it was decided to review the design and conduct another trial. This trial was also conducted on sandstone. A different chemical powder was used, one designed for 89 – 130 mm holes. The trial parameters are summarised in Table 4.

The chemical fractured the rock. The crack width stopped increasing after 96 hours when it reached a crack width of 100 mm, as shown in Figure 10.

The crack length reached saturation after 24 hours, as shown in Figure 11. This is similar to Trial 2.

The entire fracturing process is summarised in Figure 12. Figure 12(a) shows how the crack length extends from one hole to the adjacent hole. The chemical was not filled to the collar; however, it rises to the collar during the hydration process allowing the cracks

Table 3

Trial 2 parameters

Hole

Number of holes poured with the agent

Block size of poured holes

300 m2

Quantity of powder and water mixed in hole 40 kg of powder (8 bags) and 12 L of cold water

Time taken to mix and pour 1 hole 2 minutes 30 seconds

Chemical type

30 mm – 50 mm. Same chemical as in Trial 1

Depth remaining after pouring No cracks, not measured

Table 4

Trial 3 parameters

the edge/free face

Number of holes poured with the agent

Block size of poured holes

Quantity of powder and water mixed in hole

m2

80 kg of powder and 24 L cold water

Time taken to mix and pour 1 hole 3 minutes 30 seconds

Chemical type

89 mm – 130 mm hole diameters

Depth remaining after pouring No cracks, not measured

to form from the surface. Figure 12(b) and Figure 12(c) show that the crack length develops first, before the fracturing can proceed in all directions around the perimeter of the hole. This can be observed from the omnidirectional mini-cracks propagation around the

Figure 7—Crack width for Trial 2 after about 80 hours

Figure 8—Crack length for Trial 2 after approximately 80 hours

Figure 9—Fragmented rock for Trial 2

The use of soundless chemical demolition agents in large scale in situ rock breaking

Figure 10—Crack width for Trial 3 after approximately 100 hours

Figure 11—Crack length for Trial 3 after approximately 100 hours

Figure 12—Fracture results obtained from Trial 3 – (a) Cracks between adjacent holes, (b) multiple smaller cracks development, (c) fully developed cracks between two holes, and (d) final fragmentation results

holes. When the chemical still has expansion power, that power is used to widen the cracks after the crack length has developed. This shows that to achieve a wider crack, hole spacing should be kept at a minimum, so that the crack length develops faster. Figure 12(d) shows that the chemical fractured the rock to the edge.

The fractured rock was also loaded with an excavator as shown in Figure 13. The chemical was successful in fragmenting the entire depth of the hole.

The chemical was used in the correct hole diameter. The chemical has added chemical admixtures different to the chemical used in the first two trials. The admixtures prevent blow outs. The quantity and type of the chemical admixtures were not available to the public at the time of study, as this is a new product and it is still being tested.

Trial 4: Mine C

This trial was conducted on norite rock. Table 5 shows the trial parameters.

Figure 14 shows the area that was prepared for Trial 4. Three 45 m blocks were fractured separately. The block was not fractured at once. The holes that had been filled with chemical were fractured first and then moved towards the centre by a few rows at a time.

Figure 13—Fragmentation results and ease of loading the fragmented material

Table 5

Trial 4 parameters

Distance from the edge/free face

Hole burden

Number of holes poured with the agent 157

Block size of poured holes

Quantity of powder and water mixed in hole

490.9 m2, 45 m diameter block

20 kg of powder and 6 L cold water

Time taken to mix and pour 1 hole 2 minutes 30 seconds

Chemical type

89 mm – 130 mm hole diameters. Same with Trial 3

Depth remaining after pouring No cracks, not measured

The chemical fragmented the rock. The crack width stopped increasing after 72 hours when it reached a crack width of 50 mm, as shown in Figure 15.

Trial 3 had the smallest spacing when compared to the burden for all trials and the fastest growth in crack width development after 24 hours. This further proves that when spacing (crack length) is kept at a minimum, the crack width will develop at a faster rate as the remaining power is used to widen the cracks. The trial used a hole diameter that is 28% smaller than recommended, this shows that the chemical can also be used in actual hole diameters smaller than the theoretical hole diameter. The chemical also fractured the entire hole depth. Crack length was saturated after 24 hours, as shown in Figure 16.

Trial 5: Mine D

This trial was conducted on dolomite rock with the parameters given in Table 6.

The use of soundless chemical demolition agents in large scale in situ rock breaking

The trial area is shown in Figure 17, consisting of two blocks. The fragmented rock in Block 1 was fractured first to create a free face for Block 2. Block 2 was the drilled-and-filled with the product.

The chemical fractured the rock. The crack width ceased to increase after 96 hours when it reached a crack width of 75 mm, as shown in Figure 18.

The crack length was saturated after 72 hours, as compared to 24 hours in other trials, shown in Figure 19. This was due to the use of the Nex-Pand powder in a lower ambient temperature than it was designed to work in (25˚C ≤ Temp ≤ 40˚C).

Discussion and analysis of the results

The trials were compared based on rock strength, powder quantity, hole diameter, burden, spacing, and the distance to the free face. Table 7 summarises the parameters used in the comparison during the trials.

The crack length is a function of hole burden and spacing. A bigger hole burden and spacing will lead to a longer crack length. Hence, in Figure 20, Trial 3 was observed to having the highest crack length, followed by Trial 5.

Table 6

Trial 5 parameters

Number of holes poured with the agent

Quantity of powder and water mixed in hole 60 kg of powder and 18 L cold water

Time taken to mix and pour 1 hole 2 minutes 30 seconds

Chemical type

89 mm – 130 mm hole diameters. Same with Trial 3 and Trial 4

Depth remaining after pouring No cracks, not measured

Figure 21 shows the crack width for all the trials. Trial 4 was conducted on the hardest rock. It has the smallest crack width. More energy is used to develop cracks instead of widening them. Trial 4 has the same hole parameters (hole diameter, half the depth and

Figure 14—45 m diameter block being filled with the chemical

Figure 15—Crack width for Trial 4 after approximately 80 hours

Figure 16—Crack length for Trial 4 after approximately 80 hours

Figure 17—Trial 5 (1) An area with fragmentation results and (2) an area prepared for breaking

Figure 18—Crack width for Trial 5 after approximately 100 hours

The use of soundless chemical demolition agents in large scale in situ rock breaking

powder quantity, and same confinement). However, with Trial 2, the crack width is much smaller than in Trial 2. This is because of rock strength. Norite is stronger than dolomite.

A bigger quantity of Nex-Pand powder per hole generates more expanding pressures. The crack width is expected to be the greatest on holes with higher powder quantity because more pressures are produced. Trial 3 has the highest powder quantity followed by Trial 5, then Trial 2. However, Trial 2 has the greatest crack width as compared to Trial 3 and Trial 5. This is due to the confinement. Trial 2 is less confined compared to Trial 3 and Trial 5. This proves that both confinement and powder quantity affect the crack width, but confinement has the greatest influence on crack width compared to powder quantity. Trial 5 proved that the use of Nex-Pand powder in lower temperatures than designed for would result in a slower generation of cracks. The trials proved that SCDAs can be used

to fragment large volumes of in situ rock. They can be applied to fragment the rock in Block 10 at Mine A and in other areas closer to communities and sensitive infrastructure.

Comparing SCDAs to explosives

Explosives, unlike SCDAs, provide the ideal fragmentation at a faster rate, however, their use is limited in areas closer to communities and sensitive structures. Table 8 compares explosives with SCDAs based on blast design parameters and manner of breaking. When designing a blast, rules of thumb are used. The results of the five conducted trials were used to determine the rules of thumb for SCDAs.

From Table 8, the burden for explosives is significantly more than for SCDAs; this allows for a bigger area to be fragmented when using explosives. The rule of thumb used when determining the spacing for both explosives and SCDAs is the same. Explosives allow for cast blasting pattern, which reduces the handling of the fragmented rock. The rock fragmented using SCDAs remains in one place. There is more material handling with the use of SCDAs. The bench height in both SCDAs and explosives is a function of the burden used. Explosives allow for a longer bench height than in SCDAs. SCDAs do not require stemming. Timing is used in explosives to get the sequence of the blast, however, in SCDAs, the sequence of breaking is controlled by how the chemical is poured into the holes. The block area is not broken at the same time but in rows. The row poured with the chemicals first and left to break will be the first one to fragment, followed by the next row to be poured with the chemical and left to break.

Table 7

Parameters of all trials

Figure 19—Crack length for Trial 5 after approximately 100 hours

Figure 20—Crack length over time for all trials

Figure 21—Crack width over time for all trials

The use of soundless chemical demolition agents in large scale in situ rock breaking

Table 8

Explosives vs SCDA factors

Rock

Hole diameter (D)

Burden (B)

Spacing (S)

Blast pattern ratios

20 – 40 * (D)

(Number increases with a decrease in rock strength)

1 – 1.5 * (B)

Cast: S/B = 1.5

Box cut, tight breaking, and square:

S/B = 1

Staggered patterns: S/B = 1 – 1.5

Equilateral: B/S = 1 – 1.15

Bench height (H)

Stemming length (L)

Initiation/ timing

Pre-splitting

Sub-drill

Conclusion

2 – 4.5 * (B)

60 – 40 * (D)

20 – 40 * (D).

Aggregate: 2: – 30 * (D)

Drill chippings: 30 – 40 * (D)

Inter hole = 2 – 4 * (S) ms/m

Inter row = 8 – 18 * (B) ms/m

Hole spacing:

No free face: 6 – 10 * (D)

With free face at some distance:

8 – 12 * (D)

Closer free face: 15 – 20 * (D)

Hole diameter, open cast coal mines: 127 – 311 mm

0.2 – 0.5 * (B)

8 – 12 * (D)

The aim of this study was to investigate if SCDAs can be used to fragment in situ rock on a larger scale for non-explosive rock breaking applications in the mining industry. These expansive products have several advantages over explosives in that they do not cause any ground vibrations, noise, fly-rock, or harmful gases, except for their slow process. Previous studies have only been conducted in the construction and demolition industries with very limited application in the mining industry on a larger scale. In this study, a methodology was developed to conduct five field trials at four different mining sites. Results and observations showed that the chemicals, when mixed with water, expand and rise to the collar of the hole during hydration, but this will either result in blowouts or the development of cracks. When designing an area for the use of SCDAs, the hole spacing should be kept at a minimum to achieve wider cracks and yield good fragmentation results. The crack length develops first and thereafter the remaining expanding power is used

Yes

Yes

Yes

NA

NA

Yes

Small hole chemical: 28 – 64 mm large hole chemical: 64 mm – 130 mm

Large hole: 2 – 5 * (D)

– 1.5 * (B)

There is no throw of the muck pile, as the rock only fractures and remains in one place. For this reason, cast pattern is not applicable.

Mostly breaking with one free face, thus square to staggered pattern will be used.

S/B= 1 – 1.5

3 – 6 * (B)

Holes are left uncovered after pouring the mixed SCDA.

There is no initiation. The holes are poured with chemical and left to break.

No crack growth required into sidewall. Uniform vertical face required.

Effects of SCDA to be investigated.

NA

No tests are currently conducted on coal to determine the effect of sub-drill.

to widen the cracks. The powder quantity and rock confinement have the greatest influence on the crack width generated. However, confinement was observed to having the greatest influence on the generated cracks compared to powder quantity. A high powder quantity produces more expanding pressure, resulting in wider cracks. An increase in the confinement can result in much less, to no cracks, due to the tensile stress produced from the powder being less than the tensile strength of the rock. Rocks that have a higher tensile strength require a larger powder quantity and reduced hole burden and spacing to produce the same results as low strength rocks. The use of SCDAs in areas of low ambient temperature than recommended, delays the saturation of crack length. For optimal fragmentation, the SCDAs should be used as per the manufacturer’s recommended specifications, as deviations can result in slower to no cracks being generated. The trials successfully proved that SCDAs can replace explosives rock breaking method in areas that are closer to communities and sensitive infrastructure.

The use of soundless chemical demolition agents in large scale in situ rock breaking

Future studies

Further research work is being conducted to optimise the reaction rate to reduce the hydration duration in order to fragment the rock much faster than is the case currently. This will require an investigation of the interaction of the relevant chemical and physical properties. More studies should be conducted to determine ways in which the crack development can be optimised while reducing the powder quantity in the holes. Another study is suggested to investigate the effects of the presence of water in a hole on pressure generation.

Acknowledgements

Dylan Harlen (Sales Executive at Harlen Quarry Supplies): Assistance in data collection.

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The use of soundless chemical demolition agents in large scale in situ rock breaking

Affiliation:

1Department of Mining Engineering and Mine Surveying, University of Johannesburg, South Africa

2Institute of Innovation, Science and Sustainability, Federation University Australia, Australia

3Department of Metallurgy, University of Johannesburg, South Africa

4Faculte des Sciences Appliquees, Universite de Mbuji-Mayi, Democratic Republic of the Congo

5Faculty of Business and Law, University of Northampton, England

Correspondence to:

O.D. Eniowo

Email: oeniowo@uj.ac.za

Dates:

Received: 29 Dec. 2024

Revised: Mar. 2025

Accepted: 14 May 2025

Published: July 2025

How to cite:

Eniowo, O.D., Onifade, M., Grobler, H., Mulaba-Bafubiandi, A.F., Otuogbai, O.S. 2025. Evaluating the creditworthiness of a viable artisanal and small-scale mining operation.

Journal of the Southern African Institute of Mining and Metallurgy, vol. 125, no. 7, pp. 385–392

DOI ID:

https://doi.org/10.17159/2411-9717/3627/2025

ORCiD:

O.D. Eniowo

http://orcid.org/0000-0001-6173-7709

M. Onifade

http://orcid.org/0000-0001-9933-266X

H. Grobler

http://orcid.org/0000-0002-4729-5753

A.F. Mulaba-Bafubiandi

http://orcid.org/0000-0002-3437-1622

O.S. Otuogbai

http://orcid.org/0009-0000-3882-1865

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

by O.D. Eniowo1, M. Onifade2, H. Grobler1, A.F. Mulaba-Bafubiandi3,4, O.S. Otuogbai5

Abstract

Artisanal and small-scale miners in Nigeria struggle to attract formal financing to upgrade their operations to a more sustainable form of small-scale mining operation. Literature shows that for an investment to attract formal sources of financing; the business must be considered creditworthy by potential lenders. This means that the credit risks associated with the investment must be bearable for the potential lender. This study looks at how artisanal and small-scale miners can attract formal funding, which will help them take the leap from artisanal state to a form of sustainable small-scale mining operation. This study evaluates the creditworthiness of artisanal and small-scale mining operators using key credit risk parameters identified in the literature and through primary investigation. A quantitative study is conducted on a sample of 100 artisanal and small-scale miner establishments in Nigeria to evaluate the effect of these identified credit risk parameters on creditworthiness. It is shown in this study that being a member of a mining cooperative society, which is a form of social collateral, best improves the chances of access to formal financing for miners within the sample group. This underpins the role of social collateral, in both the formal and informal economy, and its influence in improving the creditworthiness of a typical artisanal and small-scale miner.

Keywords funding, loans, mining, credit, formalisation, financial sustainability, investment risk, viability assessment, artisanal and small-scale mining

Introduction

Artisanal and small-scale mining (ASM) – a low tech and labour-intensive form of mineral extraction and processing – continues to grow as an occupation across developing countries (Hilson, 2016). Available data shows that more than 30 million people globally are directly employed by the ASM industry (Stocklin-Weinberg et al., 2019). A recent estimate shows that over 2 million people in Nigeria depend on ASM directly for their livelihood (Abuh, 2023), even though the occupation has been widely associated with the environmental pollution it creates, the attributed safety and health hazards, and other social concerns (Environmental Law Institute, 2014; Malone et al., 2023; Owolabi, et al., 2017). As the population of those in ASM increased exponentially over the years, some scholars have argued that the rush to the occupation is born out of the desire to tap into the attractive minerals market in order to “get-rich-quickly” (Verbrugge, 2015; 2016), while others argue that the occupation is mainly driven by poverty, which is prevalent in several communities hosting mineral resources across sub-Saharan Africa (Hilson, 2010; Siwale, Siwale, 2017). However, recent studies such as Traoré et al. (2024) and Hilson and Hu (2022) have put this debate to bed by arguing that any attempt to dismiss the “poverty-driven” nature of ASM on the account of few individuals in it who have accumulated wealth over time, is misleading because their new-found circumstances do not disprove the reason they chose to pursue work in this sector in the first place. Thus, it has been widely agreed that a primary motivation for people in this occupation is the need to seek succour owing to the prevailing rural poverty (Hilson, Hu, 2022).

It is, therefore, essential that those in this sector are supported in a way that their operations can run in an economically, socially, and environmentally sustainable manner. Specifically, considering the economic importance of the occupation, experts have agreed that a path towards formalising the operations should be pursued rather than criminalising it (Martinez et al., 2023). The drive towards formalisation has achieved notable successes across sub-Saharan Africa. In the Nigerian ASM sector

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

for example, there are currently 251,500 registered miners and 633 small-scale mining companies (Abuh, 2023). But this has not translated into an efficient and sustainable mining operation in the sector (Eniowo, 2024). The main reason is connected to the inability of the miners to attract formal funding that could help advance or even sustain their operations. Typically, investors in the mining industry seek funding from two probable sources – formal and informal sources (Eniowo, Meyer, 2020). While artisanal and small-scale miners usually find it easier to access informal sources owing to the simple application procedures involved, they find it difficult to gain access to formal financial sources due to their bureaucratic procedures, ill supervision, and low sustainability of the mining operations (Paschal et al., 2023). In a study by Amedu, et al. (2023) conducted in the Nasarawa state of Nigeria, it was found that up to 80% of the ASM operators in this study area fund their operations with personal savings. The study further found that the remaining 20% of these miners obtained loans from friends and family members, while not a single operator has successfully funded his operation through bank loans.

Although accessible, informal financing has its own challenges as well. Across sub-Saharan Africa, there are reports of miners being exploited in these informal funding arrangements (Fold et al., 2014; Perks, 2016). This creates a poverty trap for the miners, reducing their capacity to engage in deep-level mechanisation to access abundant mineral deposits (Mallo, 2011; Paschal et al., 2023). Achieving meaningful improvement in this sector begins with capturing more and more miners in the formal domain through the existing formalisation drive. However, sustainable formalisation requires the acquisition of requisite permits that allows one to operate legally, have advanced equipment, access modern and affordable technical capacity, and most importantly, financial muscle to run their operations safely. All these factors are important for the sustainability of their operations. It has been established in recent literature that, the pillar to achieving sustainable formalisation includes exposing such miners to more knowledge on business development, access to credit facilities, and also commercialisation of the mineral proceeds and corporate governance (Martinez et al., 2023). Specifically, Amedu et al. (2023) opined that for banks to be disposed to ASM lending, there is need to establish risk-sharing mechanisms that mitigate ASM-related risks.

It is common knowledge that ASM establishments operate in poor and unsafe conditions that pose danger to health, safety, and the environment (Mushiri et al., 2017). However, studies concerned with ASM have shown that there are linkages between the dangerous and unsustainable practices by the miners and their inability to attract investment funds to upgrade the operations to a more sustainable form of operation (Eniowo, Meyer, 2020; Eniowo et al., 2022a; Seccatore et al., 2014). For example, scholars have argued that the absence of formal financing opportunities for mechanised operations or investment in acquisition of geological data to support the operations have pushed miners to adopt short-term mineral extraction strategies, which rely on informal arrangements with outside financiers, traders, and sometimes largescale mining companies, just to access the global mineral market (Perks, 2016). On the one hand, such a form of capitalisation leads to the adoption of unsafe mining techniques (Mallo, 2011), and on the other hand, it results in some form of exploitation of the miners by the financiers (Fold et al., 2014) and prevents the miners from having the decision-making power to plan far into the future regarding their business (Hilson, Ackah-Baidoo, 2011; Perks, 2016). Consequently, the report of Intergovernmental Forum on Mining,

Minerals, Metals, and Sustainable Development (IGF) (2018) asserts that the financiers, which the study described as ‘power holders’ in ASM, make the largest share of profit, while those doing the work on the ground barely make enough to survive. However, there is a scarcity of studies that develop models for enhancing the ability of ASM to attract investment capital. Rather, most studies have focused on the safety concerns and the effect of the release of harmful substances by these miners on the environment (Clement, Olaniyan, 2016; Environmental Law Institute, 2014; Mallo, 2011).

Formal lenders are usually reluctant to lend to ASM-related activities for several reasons. Some studies blame the lack of understanding of local banks on how to translate geological assets into a form of collateral with which they are familiar (Perks, 2016). Other researchers point at the inability of ASM operators to present proof of availability of ores that will assure recovery of costs and profit margins, and the lack of technical competencies, documents proving ownership, and important supporting information on the feasibility of the investment (Eniowo et al., 2022b; Marin et al., 2016; Spiegel, 2012; Van Bockstael, 2014). For these reasons, lenders consider ASM to be a high-risk investment with no guarantee of return or financial success (Marin et al., 2016). The bottom line is that beneficiaries of banks’ investment capital must be creditworthy companies with financially viable operations. For ASM firms to have access to formal sources of funding, their operations must be considered viable by potential investors and lenders. Again, for an investment to be creditworthy, the credit risks associated with the investment must be acceptable to the lender. Generally, there are several variables that lenders consider before providing loans for mining operations. For ASM, a review of the literature identifies some perceived credit risks that leads to the apathy of formal lenders in the industry. They include the assurance of availability of the mineral resource, availability of collateral security for the credit facility, availability of social security for the borrower, and the ability to pay an interest rate based on the level of productivity. Others include the miners’ technical and financial management competence and the proof of the existence and viability of the orebody to be mined (Amedu et al., 2023; Eniowo et al., 2022a; Reichel, 2019).

The existing conventional method for evaluating the viability of a typical mining operation involves a careful geological exploration research, detailed analysis, review, and modelling of technical data on the indicated resources, and proving such resources to be a viable reserve (Seccatore et al., 2014). This process is usually complex, lengthy, and expensive. Most conditions demanded by formal lenders can only be met by large-scale mining (LSM) operators. In practice, it is only when these outlined processes have been completed that a bankable feasibility study report that is pleasing to a formal lender can be produced (Rupprecht, 2004). Owing to previous failed attempts at accessing formal financing, ASM operators know that they cannot easily attract funds from formal sources of funding. Apathy in formal finance for ASM operations therefore exists on both sides of the divide, that is, amongst both formal lenders and ASM operators (Eniowo et al., 2022b). Quantification of credit risk involves assigning measurable and comparable numbers to the likelihood of default risk to loans (Ross, 2020). In literature, the variables required for analysing credit risks can broadly be classified into financial and nonfinancial variables. Most studies on credit risk prediction focus on the use of financial variables. In such models, default risks are computed based on the available data from the credit history and credit report of the loan applicant. However, for small companies, studies have shown that

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

non-financial variables particularly play key roles in influencing default risk. For example, Kanapickiene and Spicas (2019) observed that non-financial variables, such as indicators of litigation and behaviours in social networks, critically affect the default risks of small companies. In ASM, the adoption of non-financial variables for the prediction of credit risks could even be more important, owing to the rudimentary nature of the occupation where proper documentation of productivity is sometimes non-existent. Also, the unavailability of external rating agencies that can provide credit ratings for small companies in some developing countries creates a bottleneck in the use of credit reports for credit risk predictions for such operations. Thus, it is essential that a simplified method be developed to estimate the viability of ASM operations – a method that is void of the encumbrances seen in banks’ due diligence for larger mining companies. The scarcity of such, in the existing literature, necessitates this study. This paper therefore evaluates the creditworthiness for a viable small-scale mining operation using identified credit risk variables. Before going deep into this discussion, it is important to look briefly at the policy directions for the Nigerian ASM sector – the study area for this research.

Policy directions of the Nigerian ASM sector

The artisanal and smalls-scale mining (ASM) sector is considered a peculiar sector for the economic development of Nigeria. Mining of metals in Nigeria has a long history dating back to the 19th century, several decades before the promulgation of the Nigeria Minerals Act of 1946 (Mallo, 2011). Prior to the political independence of Nigeria in 1960, Nigeria was a major producer of metals, supplying more that 10% of the global output of tin concentrate (FELL, 1939). In the 1960s, the mining sector accounted for approximately four to five percent of the nation’s GDP (PWC, 2023). After Nigeria had gained its independence from Britain in 1960, many factors contributed to the collapse of her large-scale mining industry. They include the indigenisation decree of 1973, a decline in global metal prices (including tin) in the 1980s, the Nigerian crude oil boom in the 1970s, and subsequent overdependence on the oil sector. Others include the depletion of alluvial reserves, the Nigerian civil war (1966–1970), and ineffective state control (Oramah et al., 2015; World Bank, 2012). In effect, ASM now dominates the country’s solid mineral sector, representing up to 95% of the industry (Oramah et al., 2015). Consequently, the contribution of this sector to the economy of Nigeria has since dropped significantly, representing a meagre 0.17% of the GDP in the years 2018 to 2022 (PWC, 2023). However, in recent years, officials of the Nigerian government have identified this sector as the potential alternative for the needed diversification of the country’s economy, away from its dependence on petroleum sector. The current elected government, which was inaugurated in May 2023 released a policy framework to support the country’s sustainable and inclusive economic growth. This is in furtherance to the Economic Recovery Growth Plan (ERGP) developed by the immediate past government. Furthermore, in this new framework, unlocking the potential of the mining sector was identified as a major deliverable. The roadmap to unlocking the potential of this sector involves repositioning it to attract large international mining players. It also involves implementing a framework to formalise and legitimise informal ASM activities. It is noteworthy that previous governments had also identified the potential of the country’s ASM subsector. Specifically, the policy direction of the immediate past government focused on upgrading and upscaling the existing ASM industry. Based on this policy direction, the then Ministry

of Solid Minerals projected, albeit ambitiously, an increase in the GDP contribution of the mining sector from less than USD1 billion in 2016 to USD25 billion in 2026 (Akintola, 2016). The potential of the solid mineral sector is even further energised by the recent discovery of significant deposits of high-grade lithium in various parts of Nigeria, establishing the country as one of the nations that is rich in lithium resources worldwide (Amans et al., 2023). In its review of mining sector performance from 2017 – 2020 in line with its Economic Recovery Growth Plan (ERGP), the previous government stated that it had increased the number of registered mining cooperatives from 600 to 1,495. It had provided a total of 500 members of these cooperatives with extension services. It had also trained more than 1,000 of them on safer mining practices and occupational hygiene through the safer mining pilot project (Federal Government of Nigeria, 2021). Officials of the government recognised that to optimise the contribution of ASM to Nigeria’s economic development, the issue of insufficient funding must be tackled (see Figure 1). In this regard, the Minister pinpointed insufficient funding and lack of access to capital as major factors militating against ASM operations. To combat this challenge, the Ministry of Mines and Steel Development (MMSD), in collaboration with the Nigerian Bank of Industry (BOI), launched a N5 billion (then approximately USD14 million) fund in support of ASM miners in the country. The government ministry under that administration also secured other funding interventions. Some of these funds include a USD100 million fund secured from the mining sector component of the Natural Resources Development Fund sponsored by the Nigeria Federal Government and USD150 million from the World Bank for the Mineral Sector Support for Economic Diversification (MinDiver) programme (Mohammad, 2017). How these various interventions impacted the output of ASM operations is the question that remains unanswered in academic and professional discussions centred on the Nigerian ASM industry. However, available data have shown that the contribution of the mining sector (excluding petroleum and natural gas) to the country’s economy is still minimal after seven years since the government launched the intervention by the BOI in 2017. The current contribution of the sector to the nation’s nominal GDP is still in the neighbourhood of USD1.5 billion, translating to less than 1% of the economy (National Bureau of Statistics, 2023). Mining activities are capital-intensive in nature, and as such, capital plays an important role in providing the needed external financing for

Figure 1—Firms ranking obstacles to business in Nigeria (Source: World Bank, 2014; Salati, 2024)

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

mining operations. A cursory look at the disbursement of the BOI loan scheme indicates that there has been a low level of access by ASM miners. At a workshop with miners in December 2020, more than three years after the scheme kicked off, the government minister expressed displeasure over low access by miners to the loan scheme: by the end of September 2020, only 138 completed applications totalling N14.59 billion (approximately USD19.45 million) had been received and were at various stages of processing. In this same period, only 13 loan applications totalling N1.08 billion (approximately USD1.44 million) were approved under the scheme. Moreover, the rate of disbursement is also low when compared with the total applications. The reason ascribed to this by the government ministry is that applicants have largely failed to meet the necessary but stringent requirements of the BOI who handled the disbursement. This adds credence to the existing argument in literature that the stringent conditions attached to formal loans usually force artisanal and small-scale miners to opt for informal loans, which are relatively quicker, involve cheaper initial transaction costs, and lack the burdensome collateral requirements (Eniowo et al., 2022a; Spiegel, 2012). Also, while the BOI fund, like other formal monetary interventions, may offer lower interest rates, the extended loan processing timeline may not fit into the planning schedule of a typical small-scale mining operation. ASM miners are not preferable clients to lenders because they are assumed to carry more credit risks. It is therefore important to briefly go through the process of credit risk estimation adopted in this study.

Methodology

The study involves two phases. The goal of the first phase is to collate the key variables that influence creditworthiness of a viable ASM operation. This phase of the study adopts a qualitative approach, which was used to collect data through primary and secondary sources. The primary source of collection involves face-to-face interviews conducted with two groups of respondents. The interviews with the first group of respondents were conducted between September 2021 and November 2021 with bank authority, to find out about the policies that guide bank loans, especially to ASM operators. A follow-up interview was conducted in October 2022. The bank was purposively selected because, unlike the commercial banks in the study area, the bank is disposed to providing loan packages for ASM operations. The secondary source of data include documents retrieved from this bank, such as a loan application form checklist, which helped in identifying the key parameters required by the bank for a competent loan application by an ASM operator. The second group of respondents in the first phase of this study involved a selected sample of ASM operators from South West Nigeria who provided insights on the factors that affect access to credit in the industry. South West Nigeria was purposely selected for the study because it hosts various types of mineral resources commonly mined through ASM operations, which include gold, gemstones, industrial minerals, sand, and laterite (Nigeria Ministry of Foreign Affairs, 2022).

The second phase of the study adopts a quantitative approach. In this phase, data were obtained from a total of 100 purposively selected ASM establishments in the southwest region of Nigeria using a well-structured questionnaire. While previous studies on credit risk modelling for small-scale companies adopt the use of historical data, the unavailability of such data on the operations of the Nigerian ASM industry necessitates the use of current data. The major criteria for selection of respondents include the level of operation (artisanal operations that use manual methods of

operation were targeted). One copy of the research questionnaire was issued to each of the sampled operations. Out of the operations sampled, only about 20% have had previous access to formal sources of loans to fund their operations. A multiple logistic regression model was then developed to predict the creditworthiness of a small-scale mining operation using the identified credit risk variables. The adoption of logistic regression follows the use of logit scoring models to predict credit default risks of small and medium-sized businesses (SMEs). This method was used in Behr and Güttler (2007), Altman and Sabato (2007), and Kanapickiene and Spicas (2019) to measure individual credit risks. In logistic regression models, the dependent variable is binary or dichotomous. The model analyses the relationship between multiple independent variables and a categorical dependent variable and estimates the probability of occurrence of an event by fitting data to a logistic curve (Park, 2013). The models are sometimes called logit models. These models can be classified into two types, binary logistic regression and multiple logistic regression.

In this study, the regression analysis was done in R studio, and the results were verified through an analysis done on SPSS software. In the multiple logistic regression, identified variables from phase one of this study (predictor variables) were regressed with access to formal fund (dependent variable) to determine the effect of the identified variables. The dependent variable is a dummy variable that takes value 1 if an ASM operator receives a formal loan and 0 otherwise. The predictor variables are the potentially relevant parameters that may drive creditworthiness. Specifically, based on the first phase of this study, seven parameters were identified to be critical in assessing creditworthiness for ASM loan applications. They include availability of mining license (and land tenure), availability of a minimum required reserve estimate to guarantee continuity of operations, ability to pay the 5% interest rate required by the BOI for ASM loans, ability to provide a guarantor, availability of a feasibility report, proof of membership of a registered mining cooperative society, and proof that the operation is brown field (the mine site must have been running successfully for at least a year). After retrieval of the completed research questionnaire from the sampled group of respondents, it was discovered from preliminary analyses that all the respondents had licenses and land tenure (expectedly, since the list of ASM establishments was accessed through the government ministry of solid minerals). Also, about 99% of the miners within the sample group have operated their mine sites for more than the required one-year period. Similarly, the variable “availability of a feasibility report” was found to be nonstatistically significant in this study, as most of the miners argued that even though they did not have this document at the time this study was conducted, they could access it through the accountant in their cooperatives who provides financial advice and documentation for loan applications. Therefore, these three parameters were kept as constant for the purpose of the regression analysis. The form of the basic logit model is shown as: [1]

where P is the probability that a small-scale mine operator will receive a formal loan; a is the coefficient of the constant term; β is a vector of coefficients of the independent variables; X is a vector of independent variables, and e is the error term that is lognormally distributed by assumption. The coefficient of the constant and the vector β are estimated through maximum likelihood estimation. The transformation of the dependent variable constrains P to be in

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

the interval [0,1]. This standardisation is one of the main advantages of logit regression models and allows for the computation of P by a borrower by simply plugging the borrower-specific variable values into the estimated logit function (Behr, Güttler, 2007).

Hypotheses

The dependent variable in the logistic regression model is the observed access to credit. The following independent (predictor) variables were used in the logistic regression analysis: availability of a minimum reserve to guaranty continuity of operation and availability of a guarantor to stand as surety for the loan sum (a form of collateral security in the absence of physical collateral). Others include the ability of the applicant to afford the bank’s interest rate, and the confirmation of the miner’s membership in a registered cooperative society (this is considered a form of social collateral). The four predictor variables are categorical and binary in nature (true/false, 0/1). The simplicity of the question-types helps the miners to be able to provide accurate responses, which are valuable for the purpose of the analysis. This also helps to improve the interpretability of each predictor in the multivariable model and reduces the impact of outliers, as seen with continuous variables. Thus, the choice of adoption of binary predictors is borne out of the need to ensure future replication of this methodology while considering the peculiarities of the artisanal and small-scale mining industry.

Minimum reserve

The concept of ‘minimum reserve and replication’ as proposed by Marin et al. (2016) and Seccatore et al. (2014) in estimation of the viability of an artisanal and small-scale mining operation is based on a main differential from large-scale mining. According to Marin et al (2016) and Seccatore et al. (2014), the attractiveness for external investment only lies in proving, in the early stages of the business, a minimum mineral reserve that can rapidly return the investment committed to upgrade an artisanal operation into a small-scale industrial one, with an attractive profit. In this study, to evaluate the effect of possessing a minimum reserve to guarantee the continuity and longevity of an operation in order to obtain access to formal credit, the following hypothesis was developed: H1 = Operations with a minimum reserve to return investment have a better chance of attracting formal credit.

Miners were asked to complete the research questionnaire, and they were guided through unstructured interview questions on whether they have adequate ore reserve to provide for profit and continuity of the operations. The result of analysis of the responses shows a p-value of 0.000554 which falls within the confidence level of 0.05, indicating that the model is statistically significant and as such, availability of minimum reserve is a good predictor of access to formal funds.

Availability of a loan guarantor (collateral security)

A bulk of existing literature on ASM pinpoints the itinerant nature of many artisanal miners (Eniowo, Meyer, 2020; Van Bockstael, 2014). This, amongst others, occasions the apprehension of lenders in relation to ASM lending. To minimise the risk of lending to the operators, lenders seek a guarantor who will indemnify the bank against any loss that may arise from such lending. As an illustration, one of the conditions listed by the BOI for ASM-related loans is the availability of a “competent” guarantor. Usually, the bank only accepts guarantors who fall within a certain social status: some

of the permitted potential guarantors include royal fathers, high ranking political office holders, and civil servants above grade 12 in the Nigerian civil service. In this study, to examine the effect of the availability of a loan guarantor on a miner’s access to formal credit, the following hypothesis was tested:

H2 = An Operator who has a competent loan guarantor has a better chance of attracting formal credit.

In the survey, the sampled ASM group were simply asked if they have a guarantor who could provide surety for their loan application. The miners were then asked to name the profile of their potential guarantors and if they have any potential guarantor who matches the status requested by the BOI. It was observed from the result that the p-value for the model is 0.00648, which falls within the confidence level of 0.05, indicating that the model is statistically significant and as such, the availability of a loan guarantor is a good predictor of access to the formal fund.

Ability to afford the interest rate

Informal non‐banking sources that supply mainly short‐term credit usually charge higher interest rates than formal banking sources (Sarma, Pais, 2008). In the case of the BOI, the formal lender, which was adopted for this study, a blanket interest rate of 5% is usually required to be paid by applicants from the small-scale mining sector. Literature has shown that such a “moderate” interest rate could potentially be costly for miners who fall within the artisanal level of operation even though a typical small-scale operator would ideally be able to afford it. As an illustration, a study by Siwale and Siwale (2017) narrate how emerald miners in Zambia were unable to afford a European Union credit scheme set at a 5.8% interest rate, which was funded by the European Investment Bank (EIB). The study further asserts that, even when the rates were further reduced, the “neediest” miners could not afford it, and only the “mediumsized” miners eventually benefitted (Siwale, Siwale, 2017). In this study, to examine the effect of a miner’s ability to pay the required interest rate for a small-scale mining loan on the miner’s access to formal credit, the following hypothesis was tested.

H3 = An Operator who could afford the bank’s interest rate on a loan sum has a better chance of attracting formal credit.

The miners were informed of the BOI’s loan requirement in terms of interest rate, and were asked if they could afford the required 5% charged by the bank, based on their current earnings. Miners who could afford to pay up to the 5% interest rate were coded as “yes” in the analysis while miners who could not afford up to that rate were coded as “no”. It was observed that the p-value for the logit model is 0.0319, which falls within the confidence level of 0.05, indicating that the model is statistically significant and as such, a miner’s ability to afford the interest rate is a good predictor of access to formal funds.

Membership of the cooperative society (social collateral)

The literature has shown how belonging to a recognised cooperative society could be considered a form of social collateral, which is an essential condition to access funding in a predominantly informal economic activity (Eniowo et al., 2022a; Postelnicu et al., 2014). Social collateral is more important in a sector such as ASM, where physical collateral, such as assets and properties that may be used to access credit, is usually scarce. This ideology had been embraced by the artisanal and small-scale mining division of the Nigeria Ministry of Mines and Steel Development. Miners who do not have the individual capacity to secure a small-scale mining lease

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

are encouraged to merge with other miners to be registered with the Ministry as a cooperative society. There are several benefits to being a member of a registered cooperative society. One, it enables such miners to acquire the financial capacity to secure a license. Such license is acquired in the name of the registered cooperative. In addition to this, members of each cooperative do contribute amongst themselves to access mining equipment such as excavators and bulldozers on lease, which may be otherwise too expensive for individual miners to acquire. Based on the foregoing, it is expected that a miner who belongs to a registered cooperative society would have a better chance to access formal sources of credit. Hence, the study tested the following hypothesis:

H4 = An operator who is a member of a registered cooperative society has a better chance of attracting formal credit.

To evaluate this parameter, each miner was asked if they belong to a cooperative society, and if such cooperative society is registered with the government’s Ministry of Solid Minerals. It was observed that the p-value for the model is 0.000554, which is lower than the confidence level of 0.05, indicating that the model is statistically significant and as such, being a member of a recognised cooperative society is a good predictor of access to formal funding.

The result of the logit model is summarised in Table 1.

Figure 2 shows the odds ratio plot for the four predictors. The points represent the odds ratios, while the error bars indicate the confidence intervals for each. The vertical dashed line at 1 represents the point of no effect, where the predictor does not influence the odds. The x-axis is shown on a logarithmic scale for a better visualisation of the odds ratio.

Discussion and concluding remarks

Artisanal and small-scale mining operations (ASM) have suffered from lack of access to formal sources of financing because primarily, formal lenders perceive the operations to be financially risky.

The situation is further compounded by the unavailability of a documented mechanism for estimating the credit risks of ASM operations. It is noteworthy that banks have specific methodologies for evaluating the viability of mining investment, which usually require detailed financial models. The unavailability of such models required by banks renders ASM out of reach of formal bank investment. Considering its peculiarities, credit risks for ASM operations need to be modelled separately from those of large-scale operations. The importance of creating a separate credit risk model for small-scale investments has been identified and tested through empirical study (Altman, Sabato, 2007; Behr, Güttler, 2007). A similar study in the mining sector that developed credit risk models, specifically for ASM operations, remains scarce in the body of literature. Thus, this study provides a simplified methodology for estimating the risks of lending to reasonable smallscale mining operation. The study identified four key variables that are critical for predicting the credit risk of a responsible smallscale mining operation. They include availability of minimum reserve to guarantee continuity of operations, ability to afford the required loan interest rate, availability of a collateral security (or a competent loan guarantor) and availability of social collateral. This methodology is designed specifically for the estimation of expected risks of lending to ASM operations based on the factors that qualify as risk criteria in a “not so formal” industry such as ASM.

The findings of this study will help the operators of ASM activities to identify the level of the creditworthiness of their operations and help lenders to identify and adequately weigh the risk of lending to ASM operations. This study therefore concludes that the path towards upgrading ASM operations to a more sustainable form of mining should involve an effort towards improving their ability to attract formal funding. It is recommended that future study direction should consider modifying, scaling, and improving available models for proving the viability of ASM operations. In this direction, considering the nature of ASM, such

Table 1

Determinants of access to credit

Evaluating the creditworthiness of a viable artisanal and small-scale mining operation

models must not be developed using the methodology used in proving the viability of large-scale mining operations, which usually take several years to construct.

It is important that the scope of ASM formalisation efforts is expanded to expose ASM operators to the credit risks inherent in their operation. They should also be trained on mechanisms with which such risks could be abased. Also, for effective formalisation drive, one key initiative that may play a sustainable role in enhancing ASM operators’ access to credit is the strengthening of associative entrepreneurship movements such as miners’ cooperatives (Eniowo, Meyer, 2020). Firstly, this form of association would help miners to accumulate enough capital with which they could engage in more viable and safer operations (Saldarriaga-Isaza et al., 2013). Additionally, due to the perceived itinerant nature of artisanal mining, such associations provide a face for artisanal miners, a governing structure, and an avenue through which miners could be held accountable for excesses in their operations.

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Affiliation:

1Mining Technical Solutions, CSIR, South Africa

2Mining Engineering and Mine Surveying, University of Johannesburg, South Africa

Correspondence to: T. Kgarume

Email: tkgarume@csir.co.za

Dates:

Received: 17 Oct. 2023

Revised: 21 Apr. 2025

Accepted: 10 June 2025

Published: July 2025

How to cite:

Kgarume, T., van Schoor, M., Mpofu, M., Grobler, H. 2025. Light detection and rangingbased georeferencing of underground mining ground-penetrating radar data. Journal of the Southern African Institute of Mining and Metallurgy, vol. 125, no. 7, pp. 393–400

DOI ID:

https://doi.org/10.17159/2411-9717/3166/2025

ORCiD:

T. Kgarume

http://orcid.org/0009-0001-2523-2196

M. van Schoor

http://orcid.org/0000-0003-2177-5967

M. Mpofu

http://orcid.org/0000-0003-2010-7596

H. Grobler

http://orcid.org/0000-0002-4729-5753

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

by T. Kgarume1, M. van Schoor1, M. Mpofu1, H. Grobler2

Abstract

The South African mining industry has committed to achieving a state of zero harm for its workforce, with a strong emphasis on worker health and safety. Among the major safety concerns are falls of ground, a leading cause of injuries and fatalities. Ground-penetrating radar, a non-destructive geophysical method, is recognised for its ability to image structures, fractures, and geological features within the rock mass. However, ground penetrating radar data is often acquired in local coordinates, posing challenges for visualisation in mine computer-aided design or three-dimensional visualisation software. This study explores the pivotal role of light detection and ranging data in transforming ground penetrating radar data from local survey coordinates to absolute mine coordinates. A comprehensive georeferencing methodology is presented, providing the stepwise progression from the initial georeferencing of ground penetrating radar data to the ultimate integration of ground penetrating radar and light detection and ranging datasets, resulting in the creation of a ground penetrating radar-light detection and ranging three-dimensional model. The proposed approach not only facilitates the integration of but also offers a practical means of visualising the integrated datasets within commonly used computer-aided design or three-dimensional visualisation software. An essential aspect of this integration is the adoption of non-proprietary data formats, specifically American Standard Code for Information Interchange text files, ensuring broader accessibility and compatibility. The potential for integrating diverse datasets to construct insightful models of the underground mining environment is illustrated. Integration of different datasets has the potential to offer a holistic understanding of the mining environment, providing essential information to decision-makers.

Keywords

georeferencing, light detection and ranging (LiDAR), ground penetrating radar (GPR), data integration, visualisation, health and safety

Introduction

The South African mining sector is dedicatedly striving to attain a state of zero-harm for its workforce, focusing particularly on the health and safety of its workers. To effectively comprehend the risks associated with falls-of-ground (FOG), it is essential to gain deeper insights into excavation stability through the utilisation of various technologies. In 2015, the South African Mining Extraction Research, Development, and Innovation (SAMERDI) strategy was formally adopted and served as an input document to the government-driven Mining Phakisa initiative that was held in November 2015 (Singh, 2017). SAMERDI has a primary focus on achieving three main objectives: zero harm, increased operational efficiency, and reduced costs within the mining sector. One of the critical areas of emphasis identified within SAMERDI is the Advanced Orebody Knowledge (AOK) programme, which is aimed at providing mine planners, rock engineers, geologists, and other decision-makers with valuable information and knowledge that is intended to contribute to the overarching objectives of achieving both optimal extraction and zero harm within the mining operation (Mandela Mining Precinct, 2023). The AOK programme aligns with SAMERDI's zero harm objective by conducting research into technologies capable of delineating the orebody and associated geological disturbances. This research aims to provide a deeper understanding of the rock mass and to assist in mine planning and the design of rock engineering solutions. Geophysics has previously played a pivotal role in the mining industry, primarily in secondary exploration and surveying. Its application has been instrumental in gaining insights into the stability of the rock mass within areas of geological, mining, and rock engineering disciplines. Amongst the technologies that have garnered attention for understanding the rock mass stability by

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

imaging structures, fractures, domes, potholes, and similar features, is ground penetrating radar (GPR), a non-destructive geophysical method. The local mining industry has previously leveraged GPR technology. In recent years, GPR has been used in underground mining for mapping geological discontinuities (such as parting planes, faults, dome, potholes, dykes, and fractures) and imaging of orebodies within the rock mass surrounding the mining excavations (White et al., 1999; Van Schoor et al., 2006; Vogt et al., 2005). Grodner, (2001) used GPR to quantify changes in the fracture pattern resulting from preconditioning ahead of a mining face in a deep-level Witwatersrand Basin gold mine. An in-depth discussion of its historical applications is discussed by Kgarume et al. (2019).

GPR data is obtained by moving a transmitter and receiver antenna along the surface of the rock mass to produce output results/images, which are subsequently analysed to gain insights into structures that could potentially compromise the stability of the rock mass. Typically, these results are acquired in local coordinates, which correspond to the coordinates relative to some arbitrary reference point specified for the surveyed area. Consequently, these results are not inherently linked to the mine’s underground coordinate system, making it challenging to visualise and interpret the data in the mine's computer-aided design (CAD) software or other three-dimensional (3-D) visualisation software. To overcome this, light detection and ranging (LiDAR) point cloud datasets, which were collected at the same underground site, were used to georeference the 3-D GPR data. GPR and LiDAR are presently distinct stand-alone technologies, each with its own set of data acquisition and quality control (QC) requirements. Consequently, it becomes important to demonstrate an optimised data acquisition methodology for both scanning technologies and for ultimate integration of both datasets. By georeferencing GPR data with the use of LiDAR point cloud data, it becomes possible to visualise GPR data within CAD and other 3-D visualisation software using the absolute mine coordinate system.

Data acquisition

An underground survey area was identified at Royal Bafokeng's Maseve platinum mine. The mine employs the bord and pillar mining method, focusing on the extraction at the Merensky Reef within the Bushveld complex. To access the mining zones, the operation utilises a decline system. Figure 1 shows a zoomed-in mine plan of the surveyed area (yellow box).

Survey area preparation

The primary objective of the survey preparation was to streamline the georeferencing process of the GPR data by harnessing the potential of LiDAR point cloud data. While previous research by Bubeck et al. (2011) successfully integrated GPR and terrestrial LiDAR datasets to create 3-D virtual models of geological outcrops, the unique challenge addressed in this study was not encountered in their work. Their investigations were conducted on the Earth's surface, where access to Global Positioning Systems (GPS) ensured accurate georeferencing of the datasets. In contrast, the underground mining environment presents a distinct challenge due to the absence of GPS signals. To facilitate the georeferencing process, a site visit was conducted for inspection and familiarisation, involving precise measurements of the specified survey area's dimensions. Subsequently, a 3-D GPR survey was designed, covering the hanging wall of the survey area from the back area to the mining face. Survey lines were marked along the hanging wall using spray paint, with reference to four strategically

positioned roof bolts (C1, C2, C3, C4) defining the corners of the survey grid, as depicted in Figure 2. This marking strategy ensures precise referencing of the survey line positions and their subsequent identification within the LiDAR point cloud data.

Figure 3 shows the survey grid on the local project coordinate system, where the x-axis corresponds to the spacing between the survey lines, while the y-axis corresponds to the survey line distance. The positions of the roof bolts are indicated by the reference points C1, C2, C3, and C4.

3-D GPR data acquisition

The acquisition of the 3-D GPR dataset was carried out by the team from the Council for Scientific and Industrial Research (CSIR), by following the marked survey layout. This acquisition process entailed the systematic acquisition of parallel GPR survey lines that ran along the surface of the hanging wall. To ensure comprehensive coverage of the survey area, the team maintained consistent

Figure 1—Zoomed-in mine plan of the identified survey area

Figure 2—Roof bolts used as a reference point of the survey grid

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

Figure 4—Acquisition of GPR data along the marked hanging wall grid intervals between each successive survey line. The GPR data was acquired by using a RockNoggin 1 GHz GPR system, manufactured by Sensors & Software (Sensors & Software, 2023). During data collection along the hanging wall area, a Smart Handle attachment was employed to effectively deploy the GPR antenna along the hanging wall surface, as depicted in Figure 4, ensuring precise data acquisition.

The acquisition of the 3-D GPR survey data followed a systematic approach, starting from one corner of the predefined grid. For example, in Figure 3, the initial profile was acquired, spanning from point C1 to C2. Subsequently, the operator

incrementally moved along the x-axis, acquiring consecutive profiles parallel to the first line while maintaining a consistent line spacing of 0.1 metres. This sequential process continued until all the profiles, as illustrated in Figure 4, were successfully acquired. Following the acquisition of data, processing was applied to minimise noise and enhance features, facilitating easier interpretation. Table 1 provides a list of the processing steps applied, along with the description of the effect of each step on the data. Figure 5 shows one of the survey lines acquired at the site. The x-axis represents the distance along the survey line while the y-axis represents the two-way travel time of the radar pulse (i.e., the time it takes for the radar pulse to travel from the transmitter antenna to a subsurface target and back to the receiver antenna). Using the radio wave speed in rock of 0.1 m/ns, the y-axis is converted to the depth within the hanging wall. The interpreted data demonstrates that GPR can be valuable in identifying structures within the hanging wall that could negatively impact the stability of the rock mass. This, for example, includes detecting scenarios where flat-lying and dipping fractures interact, leading to the potential formation of key blocks within the hanging wall.

Using these profiles, a 3-D GPR model of the hanging wall was constructed by interpolating between the lines to create a 3-D volumetric data representation. Figure 6 shows the generated 3-D GPR model.

LiDAR data acquisition

For this task, the team from the University of Johannesburg (UJ) acquired the data using a LiDAR system, as shown in Figure 7. The survey process involved setting up the LiDAR sensor on a tripod and strategically positioning the system at various positions within

Figure 3—Survey grid on local project coordinate system

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

Table 1

Data processing steps

Processing step

Effect on the data

Move start time correction. Moves the radargram to zero start time.

Subtract-mean (Dewow). Eliminates a low-frequency component from the signal, stemming from either inductive phenomena or antenna characteristics.

Automatic Gain Control (AGC). Brings out reflectors that occur later in time on the radargram and weak because of signal attenuation. Time cut. Cuts the length of the time axis to 40 nanoseconds, as the most valuable time, and improves the visualisation of the 3-D model.

Diffraction stack migration.

Traces back the reflection and diffraction energy to their "source".

the survey area. At each position, the LiDAR sensor performed scans of the surroundings, capturing data in all three dimensions.

By combining the different scans, it is then possible to create a 3-D point cloud of the survey area. Figure 8 displays the LiDAR point cloud representation of the surveyed underground area, while Figure 9 shows a zoomed-in section of the survey area. The point cloud data is colourised using the red, green and blue (RGB) colour scale. The RGB colour scale contextualises the point cloud as it looks more like a 3-D photograph of the survey area. The survey grid is also visible on the RGB image of the point cloud.

A fundamental aspect of the LiDAR data acquisition process is the inclusion of several "known" control or reference points within the scans. These reference points play a critical role in the subsequent georeferencing of the LiDAR point cloud data. Ideally, the acquisition strategy should encompass a minimum of

Figure 6—3-D-GPR model in local survey coordinates

Figure 7—Acquisition of LiDAR data

Figure 8—RGB LiDAR point cloud data of the underground excavation (plan view)

Figure 5—Acquired GPR survey line (left) and the interpretation of the data (right)

Dipping fracture

Dipping fracture

fracture

Roofbolt fracture

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

three known control points, which could consist of surveyed and georeferenced elements like survey pegs or roof bolts. These control points are carefully positioned based on their known and precisely surveyed coordinates relative to nearby survey pegs or other fixed reference points.

Data georeferencing

LiDAR data, similar to GPR data, is initially acquired in a local or relative coordinate system. The crucial link between these local coordinates and the known absolute coordinates of control points forms the basis for converting the relative point cloud coordinates into absolute coordinates, aligned with the mine's coordinate system. By utilising georeferenced LiDAR data, it was possible to determine the absolute coordinates of GPR reference points, such as roof bolts defining the corners of the GPR grid. This information allowed for the extraction of the start and end coordinates of individual GPR lines, facilitating the transformation from the local project coordinate system to the absolute mine coordinate system. For example, the start and end positions of a line in the project coordinate system were updated from XStart = 0, XEnd = 6.5 m to XStart = 9074.5497, XEnd = 9081.0497 for the x-coordinate in the absolute mine coordinate system. Georeferencing was

accomplished by employing these start and end coordinates to interpolate the positions of each individual radargram trace and updating the trace header files with this new position information, effectively georeferencing the GPR data. Figure 10 provides a visual representation of the sequential steps involved in achieving this georeferencing process.

Figure 11 shows selected georeferenced GPR lines plotted on a mine plan of the surveyed panel and Figure 12 shows the 3-D GPR model of the hanging wall in the absolute mine coordinate system. The successful georeferencing of the GPR lines enabled the integration of the GPR data with the LiDAR point cloud data.

Integration of 3-D GPR with LiDAR point cloud data

The two datasets were integrated and simultaneously visualised to create a GPR-LiDAR 3-D model of the surveyed site. The visualisation was done in CloudCompare (CloudCompare, 2023), a 3-D point cloud visualisation software. Figure 13 illustrates the resulting GPR-LiDAR 3-D model, which can be rotated and viewed from different perspectives. Additionally, Figure 14 provides an alternative viewpoint, showing the model from the sidewall and face area, offering a detailed representation of the surveyed site. This integration of GPR and LiDAR data enhances the visualisation and

Figure 9—Zoomed-in LiDAR point cloud plan view of the survey panel hangingwall (left) and a cross-section view of the panel looking towards the face area (right)

Figure 10—Methodology followed for the georeferencing of GPR data using LiDAR data

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

understanding of the underground excavation, providing valuable insights for geological and rock engineering analysis.

To enhance the visualisation of the GPR data, the flat fracture identified in Figure 5 was traced along the different radargrams across the 3-D data. This allowed for the visualisation of the fracture as a 3-D surface horizon across the data. Figure 15 shows the surface horizon visualised along with some selected radargrams and the point cloud data.

Figure 15 highlights the importance of acquiring and integrating 3-D datasets. By interpreting surface horizons within the 3-D GPR data, it becomes possible to generate detailed visualisations that significantly enhance the ease of data interpretation.

Discussion of results

As illustrated previously, GPR data can offer insights into the stability of the rock mass, including details about jointing, fracturing, dome and pothole structures, as well as parting planes. For instance, through data interpretation, it becomes possible to detect fracture interactions that may lead to the potential development of key blocks within the hanging wall (such as in Figure 5). Identification of such structures can provide valuable information regarding hanging wall support requirements. The georeferencing process played a pivotal role in this study, bridging the gap between GPR data in the local coordinate system and the absolute mine coordinate system. The effort to integrate these datasets was met with challenges stemming from the different data acquisition systems, software packages, and the proprietary formats of the data. These complexities prompted an exploration into alternative data formats that would streamline the data integration process. The GPR data files are in a proprietary binary format, known as ".DT1," specific to Sensor & Software GPR systems, accompanied by an associated ".HD" file containing essential header details. These header details included critical information about the survey acquisition settings, such as acquisition date, trace count, point count per trace, survey start and end positions, nominal frequency, and more. Meanwhile, the LiDAR data files were provided in the ".e57" format, a recognised point cloud data format. To ensure seamless import into various software platforms and ease of data handling, the study opted for the using of American Standard Code for Information Interchange (ASCII) text files for the models. This choice not only enhanced user-friendliness but also improved data file manageability in terms of size. Figure 15

Figure 14—GPR-LiDAR 3-D model viewed from the sidewall and face area

Figure 15—Flat fracture surface horizon

Figure 11—Georeferenced GPR survey lines plotted on a mine plan

Figure 12—3-D GPR model of the hanging wall in the absolute mine coordinate system

Figure 13—GPR-LiDAR 3-D model viewed from the back area towards the face area

Light detection and ranging-based georeferencing of underground mining ground-penetrating

showcases an example of a GPR data file exported in ASCII text format, comprising columns for X, Y, and Z coordinates, alongside a scalar field representing signal strength amplitude. Figure 16 also illustrates a LiDAR data file in ASCII text format, with columns for X, Y, and Z coordinates, as well as additional scalar fields indicating properties like RGB values. These format choices were instrumental in overcoming data integration challenges and ultimately contributed to a more accessible and unified dataset for visualisation.

Figure 17 gives a summary of the process followed for the preparation, exporting, and importing of the different datasets in order to visualise the GPR-LiDAR 3-D model.

These results demonstrate the integration of LiDAR and GPR datasets for creating an enhanced model of the underground excavation for a better understanding of the rock mass. As demonstrated by Figure 15, interpreting surface horizons within the 3-D GPR data, it becomes possible to generate detailed visualisations that significantly enhance the ease and accuracy of data interpretation. Integration of GPR with other types of datasets has the potential to enhance the accuracy, completeness, and usefulness of underground information. GPR data, when combined with other datasets like LiDAR, geological surveys, or geophysical data, provides a more comprehensive view of underground structures and conditions. This comprehensive understanding, in turn, supports better decision-making, improved safety, and more efficient and cost-effective mining.

The primary motivation for transitioning from a local coordinate system to an absolute coordinate framework is to address limitations inherent in traditional geophysical data processing. While local coordinates can be effective for immediate data interpretation, the mining industry has identified a critical need to convert geophysical outputs to absolute mine coordinates to enhance data reliability and facilitate site revisitation over extended periods. Absolute coordinates provide a consistent spatial reference, mitigating the risk of losing or displacing local reference markers such as physical objects or spray paint, which are prone to degradation or removal over time.

Furthermore, integrating 3-D GPR data with georeferenced LiDAR point cloud datasets significantly improves spatial understanding of excavation geometries. This approach enables comprehensive visualisation of excavation such as the geometry, support element locations, variations in hanging wall and footwall

topography, and the complex three-dimensional characteristics of features within the immediate hanging wall.

It is also important to note that structural features, such as parting planes or fractures, may exhibit orientations that are not parallel to the hanging wall surface, and a 3-D methodology is more effective in delineating the dip and extent of such features compared to isolated 2-D survey lines.

Conclusions

and recommendations

The use of LiDAR data proved to be instrumental in the successful georeferencing of GPR data, facilitating its transformation from local survey coordinates to the absolute mine coordinates. The study provided a comprehensive georeferencing methodology, outlining a systematic step-by-step process, starting with the initial georeferencing of GPR data to the integration of GPR and LiDAR data to form a GPR-LiDAR 3-D model. This approach not only streamlined the visualisation of the integrated datasets but also promoted compatibility by using non-proprietary ASCII text files. Through this study, the research showcased the immense potential of integrating diverse datasets, thereby paving the way for the creation of more insightful and comprehensive models of the underground mining environment to enhance decision-making and contribute to the industry’s zero-harm objective.

Figure 16—Example of an ASCII file with GPR data (left) and an example of an ASCII file with LiDAR point cloud data

Figure 17—Overview of the data preparation and export process for import and visualisation in CAD and 3-D visualisation software

Light detection and ranging-based georeferencing of underground mining ground-penetrating radar data

Further enhancements to the underground georeferencing methodology and integration of LiDAR and geophysical data sets such as GPR should be pursued through ongoing research. Specific areas of potential improvement include the logistics relating to the acquisition of LiDAR and GPR data and also possible further streamlining of the georeferencing and data integration steps.

Acknowledgements

The research team thanks the Mandela Mining Precinct through the Advanced Orebody Knowledge programme for funding of the research and Royal Bafokeng Maseve mine for providing the underground site.

References

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African Pegmatite

Allied Furnace Consultants

AMIRA International Africa (Pty) Ltd

Anglo American Platinum Corporation

Anglogold Ashanti Ltd

Anton Paar Southern Africa

Arcus Gibb (Pty) Ltd

Becker Mining (Pty) Ltd

Bluhm Burton Engineering Pty Ltd

Buraaq mining Services (Pty) Ltd

Caledonia Mining South Africa

CARBOCRAFT (Pty) Ltd

Castle Lead Works

CIGroup ACE Pty Ltd

DDP Specialty Products South Africa (Pty) Ltd

De-Tect Unit Inspection (Pty) Ltd

Digby Wells and Associates

EHL Consulting Engineers (Pty) Ltd

Elbroc Mining Products (Pty) Ltd

EPIROC South Africa (Pty) Ltd

Ex Mente Technologies (Pty) Ltd

Exxaro Resources Limited

FLSmidth Minerals (Pty) Ltd

G H H Mining Machines (Pty) Ltd

Geobrugg Southern Africa (Pty) Ltd

Glencore

Gravitas Minerals (Pty) Ltd

Hatch (Pty) Ltd

Herrenknecht AG

Impala Platinum Holdings Limited

IMS Engineering (Pty) Ltd

Ingwenya Mineral Processing

Ivanhoe Mines SA

Malvern Panalytical (Pty) Ltd

Maptek (Pty) Ltd

Mech-Industries (Pty) Ltd

Micromine Africa (Pty) Ltd

Minearc South Africa (Pty) Ltd

Minerals Council of South Africa

MineRP Holding (Pty) Ltd

Mining Projection Concepts (Pty) Ltd

Mintek

MLB Investments CC

Modular Mining Systems Africa (Pty) Ltd

Murray & Roberts Cementation (Pty) Ltd

OPTRON

Paterson & Cooke Consulting Engineers (Pty) Ltd

Pump and Abrasion Technologies (Pty) Ltd

Redpath Mining (South Africa) (Pty) Ltd

Rosond (Pty) Ltd

Roytec Global (Pty) Ltd

Rustenburg Platinum Mines Limited - Union

Salene Mining (Pty) Ltd

Schauenburg (Pty) Ltd

Sebotka (Pty) Ltd

SENET (Pty) Ltd

Sibanye Gold Limited

Sound Mining Solution (Pty) Ltd

SRK Consulting SA (Pty) Ltd

StageFright Edutainment

Tomra (Pty) Ltd

Trans-Caledon Tunnel Authority

Ukwazi Mining Solutions (Pty) Ltd

VBKOM Consulting Engineers

Weir Minerals Africa

ZUTARI (Pty) Ltd

XIX 2026 International Society for Mine Surveying Congress

VENUE: CENTURY CITY, CAPE TOWN

22-24 SEPTEMBER 2026 – CONFERENCE

25 SEPTEMBER 2026 – TECHNICAL VISITS

Hosted by Institute of Mine Surveyors of Southern Africa and the International Society of Mine Surveying.

ABOUT THE CONGRESS

The International Society of Mine Surveyors (ISM) holds its congress every three years, uniting global mine surveying professionals. South Africa will host the XIX ISM Congress in 2026 in Cape Town, focusing on all aspects of mine surveying.

What is Mine Surveying?

A specialised field within mining science, mine surveying involves measurements, calculations, and mapping throughout the mining lifecycle, including:

• Planning and controlling mine operations for safety and efficiency.

• Evaluating mineral reserves and economic viability.

• Managing mineral rights and mining cartography.

• Assessing mining impacts on land and geology.

• Supporting environmental and rehabilitation efforts.

CONGRESS OBJECTIVES AND FOCUS AREAS

The 2026 Congress will showcase ISM’s six commissions and feature key topics such as:

• Mineral & Geology Studies – Understanding deposit structure and characteristics.

• Resource Assessment & Economics – Evaluating reserves and feasibility.

• Mineral Property Management – Handling acquisitions, sales, and leases.

• Mine Operations – Optimising planning and control.

• Rock & Ground Movements – Studying subsidence and mitigation.

• Environmental Rehabilitation – Ensuring responsible land restoration.

This global event will foster collaboration, innovation, and knowledge sharing, advancing the mine surveying profession.

ECSA Validated CPD Activity, Credits = 0.1 points per hour attended.SAGC Validated Activity.

Gugu

Charlie, Conference coordinator

FOR FURTHER INFORMATION, CONTACT: E-mail: gugu@saimm.co.za Tel: +27 11 538 0238, Web: www.saimm.co.za

Empowering a future-fit mineral processing industry

IMPC

IMPC 2026 will be hosted by the Southern African Institute of Mining and Metallurgy (SAIMM). The SAIMM has been in existence for 130 years, having been established in 1894 as a ‘learned society’ to support mining and metallurgical professionals during the emergence and growth of the early South African minerals industry.

Mining is of great importance to Africa in general, and particularly to Southern Africa. Africa accounts for a major portion of the world’s mineral reserves and more than half of gold, platinum group metals, cobalt and diamonds. Southern Africa produces over two-thirds of Africa’s mineral exports by value.

CAPE TOWN INTERNATIONAL CONVENTION CENTRE

IMPC 2026 will be hosted at the Cape Town International Convention Centre (CTICC). Since the inception of the CTICC in 2003, Cape Town has been proudly the number one destination for conferences in Africa, according to the latest International Congress and Convention Association (ICCA) statistics.

Cape Town, the “Mother City”, is the oldest city in South Africa and has a cultural heritage spanning more than 300 years. Cape Town is a modern, cosmopolitan city and is often rated as one of the premier world holiday destinations. The city has a large range of hotels & guest houses and modern transport infrastructure. The city has numerous activities & attractions, including Table Mountain, Robben Island, Cape Point, the Castle, V&A Waterfront, world class beaches, wine farms, nature reserves, scenic drives, hiking, whale watching, shark cage diving and fine dining.

Photo courtesy CTICC

TOWN

The SouThern AfricAn inSTiTuTe of Mining And MeTAllurgy founded 1894