Sodium Oxide-Fluxed Aluminothermic Reduction of Manganese Ore with Synergistic Effects of C and Si Reductants: SEM Study and Phase Stability Calculations

Abstract

Aluminothermic reduction is an alternative processing route for the circular economy because Al is produced electrochemically in the Hall–Héroult process with minimal CO₂ emissions if the electricity input is sourced from non-fossil fuel energy sources. This circular processing option attracts increased research attention in the aluminothermic production of manganese and silicon alloys. The Al₂O₃ product must be recycled through hydrometallurgical processing, with leaching as the first step. Recent work has shown that the NaAlO₂ compound is easily leached in water. In this work, a suitable slag formulation is applied in the aluminothermic reduction of manganese ore to form a Na₂O-based slag of high Al₂O₃ solubility to effect good alloy–slag separation. The synergistic effect of carbon and silicon reductants with aluminium is illustrated and compared to the test result with only carbon reductant. The addition of small amounts of carbon reductant to MnO₂-containing ore ensures rapid pre-reduction to MnO, facilitating aluminothermic reduction. At 1350 °C, a loosely sintered mass formed when carbon was added alone. The alloy and slag chemical analyses are compared to the thermochemistry predicted phase chemistry. The alloy consists of 66% Mn, 22–28% Fe, 2–9% Si, 0.4–1.4% Al, and 2.2–3.5% C. The higher %Si alloy is formed by adding Si metal. Although the product slag has a higher Al₂O₃ content (52–55% Al₂O₃) compared to the target slag (39% Al₂O₃), the fluidity of the slags appears sufficient for good alloy separation.

1. Introduction

Beketov pioneered the aluminothermic reduction process in 1859 [1]. Once lower-cost aluminium became available in the 19th century, the practical application of aluminothermic reduction expanded through the work of Beketov and Goldschmidt, respectively [1]. Thermic reduction with Al or Si as reductant is a well-established method of producing low-carbon ferroalloys of manganese and chromium from a molten slag [2]. Medium–carbon ferromanganese (MCFeMn) can be produced by thermic reduction or via oxygen refining of high-carbon ferromanganese (HCFeMn). Silicothermic reduction of manganese oxide-based slags is performed in an electric furnace because the exothermic heat generated by silicothermic reduction is insufficient to meet overall process heat requirements [2]. Low-carbon ferromanganese (LCFeMn) is only produced with silicothermic reduction because of excessive manganese losses in oxygen refining.

Table 1. Ferromanganese alloy compositions (mass%); n.s. = not specified [3,5,6].

	%Mn	%Si	%C	%P	%S	%Al
HCFeMn (A)	78–82	<1.2	<7.5	<0.35	<0.02	n.s.
MCFeMn (A)	80–85	<1.5	<1.5	<0.30	<0.02	n.s.
LCFeMn (B)	80–85	5.0–7.0	<0.75	<0.30	<0.02	n.s.
AMS	60–70	5–15	0.4–2.0	<0.05	n.s.	8–20
FMA	40–80	<2.5	<1.5	<0.3	<0.03	12–16
FAMS	22–24	10–15	1–2	<0.06	n.s.	10–14

presents the chemical compositions of the various carbon ferromanganese alloys [3]. The silicon reductant is sourced as silicomanganese alloy of 16–30% Si for medium-carbon ferro-manganese production, and 32–33%Si for low-carbon ferro-manganese production. The silico-manganese minimum %Si content is selected based on the target %C in the product alloy because the carbon solubility in Fe–Mn–Si–C alloys decreases drastically with increased %Si [2]. The need for low-carbon and low-phosphorus manganese alloys is motivated in the production of transformation-induced plasticity (TRIP) and twinning-induced plasticity (TWIP) steels since these grades contain high manganese (15–25% Mn), silicon (3%), and aluminium (3%) and low carbon levels (0.01% C) [4]. Furthermore, the application of aluminium-silicon-manganese (AMS), ferromanganese-aluminium (FMA), and ferrosiliconmanganese–aluminium (FAMS) complex alloys in conventional steelmaking plants is more efficient than adding individual deoxidiser alloys [5,6]. These alloys are produced in electric furnaces operated at high temperatures, using coke as a reductant to reduce Al₂O₃ to Al [6].

The application of aluminium as a reductant opens the possibility of a circular process with reduced CO₂ emissions as proposed in recent studies [7,8,9,10]. These studies explain that since aluminium is produced electrochemically in the Hall–Héroult process, and the slag and aluminium are reacted in an electric furnace, most of the process CO₂ formation can be negated. Recent work applied discarded smelter slag as the MnO source in aluminothermic reduction since much MnO slag is stockpiled [7,8]. However, these slags are currently produced as a by-product or discard product of carbothermic reduction processes. Therefore, applying aluminothermic reduction with the ore as the starting material in a single step can provide a simplified circular process [7,8,9,10]. Some examples of this approach are already being used industrially. For example, low-carbon ferro-manganese production from roasted low-grade manganese ore is described in [11]. Pre-roasting of the ore is required to convert MnO₂ to Mn₂O₃ or Mn₃O₄, thereby decreasing the violent heat release during aluminothermic reduction with MnO₂ [11]. The ore feed requirement is a minimum of 46% Mn and a Mn/Fe ratio of 9, with a P content below 0.12%. CaF₂ is added as a flux [11]. The product alloy composition ranged from 70–84% Mn, 13–16% Fe, 1–10% Si, <0.03% Al, and <0.24% P, and the metal yield was 58–64%.

Researching the fundamentals of the aluminothermic reduction reaction is complicated by the rapid release of heat from the exothermic reaction between Al and oxidants. The implication is that the reaction system is reacted under non-isothermal conditions, which complicates the kinetic analysis [12]. Oxidants other than air may be added deliberately as an additive, or the oxides targeted for the reduction reaction may form the dual purpose of oxidant and metal/alloy product former. In some studies the mixture is reacted under a protective gas atmosphere, such as Argon gas, to eliminate oxygen in the air from reacting with the feed mixture [8,9,13]. However, it has been shown that some air is required to initiate the reaction in MnO₂–Al mixtures since this mixture did not act under Argon gas [12]. Therefore, varied reaction mixtures have been reported in the literature with different particle sizes, aluminium-to-oxide ratios, mixture compaction extents, oxide and metal solution phases formed, reactor materials, and, most importantly, inert or oxygen-containing atmosphere selection, with or without ignition chemicals or oxidants [8,9,11,12,13,14]. Industrial applications typically do not use expensive inert atmospheric gases for shielding against air; therefore, process feasibility must be tested in an air atmosphere first. This effect is observed in the work of Sarangi et al. [12], where a fundamental kinetic study on MnO₂–Al mixtures was conducted in static air, following the failure of the reaction under an Argon atmosphere. Activation energy values ranging from −518 kJ/mol to −910 kJ/mol were reported for heating rates of 5 K/min and 15 K/min [12]. In comparison, 284 kJ/mol was reported for 5 µm Al reacted, versus 416 kJ/mol for nano-Al reacted at a heating rate of 10 K/min [15]. In the low-temperature (750–850 °C) reaction of MnO and Mn₃O₄ with a stirred aluminium bath in air, activation energy values of 330 kJ/mol and 36 kJ/mol were reported [16]. It is clear that the rate constant values from each study result in apparent activation energy values peculiar to the specific factors of the reactant mixture, such as particle size, mixture compaction density, and phase chemistry changes with reaction progress [12,15,16]. Thus, in the aluminothermic reduction process, the attainment of liquid slag and alloy from the aluminothermy heat and mixture formulation is the first process requirement to ensure optimal separation of the produced alloy from the slag.

Since Al₂O₃ has a high melting point, the slag chemistry must be designed to flux most Al₂O₃. Thus, the slag volume can be excessive, and the phase chemistry of the slag will influence the ease of subsequent processing for hydrometallurgical recovery of Al₂O₃ [7,8,17,18]. Since CaO and Al₂O₃ form low liquidus compounds and CaO in the slag increases the activity of MnO in the slag to promote aluminothermic reduction of MnO, the aim slag of CaO-Al₂O₃ was applied [7,8]. The leaching of Al from CaO-Al₂O₃ slag has been demonstrated in a 10% Na₂CO₃ solution, reacted at 30–45 °C [17]. Only a small quantity of the silicate phase was not leachable. Another approach is a Na₂O-Al₂O₃-based slag since the NaAlO₂ compound is easily leached in water [18]. Adequate leaching conditions were identified at a solid–liquid ratio of 1:10 with a leaching time of 60 min at 60 °C. The leachate contained Na⁺:Al³⁺ at 1.43:1. Carbonation with CO₂ gas bubbled through the solution at 80 °C resulted in a pH of 12.5–10.5 for 96.5% Al recovery as α-Al(OH)₃ (Bayerite) particles of 90% +44 μm and 98.8% alumina grade, suitable for feed to the Hall–Héroult process after calcination to α-Al₂O₃ [18].

This work aims to formulate and test a suitable slag formulation for the application of circular processing of aluminium as the primary reductant in manganese ore reduction, producing a low-carbon ferro–manganese alloy. The slag must have high Al₂O₃ solubility to effect adequate alloy–slag separation and must allow the Al₂O₃ product to be recovered via hydrometallurgy. The synergistic effect of carbon and silicon reductants with aluminium is illustrated and compared to distinct scenarios, as well as to the test result with only carbon as a reductant.

2. Materials and Methods

2.1. Materials

The reaction mixture consisted of manganese ore, coal, metal powders of aluminium and silicon, and lime and sodium silicate as flux. The chemically pure metal powders were sourced as follows: Al (99.7% Al, −1 mm) and Si (99.9% Si, −44 µm) provided by Sigma-Aldrich (Johannesburg, South Africa). The chemically pure sodium silicate is supplied by Sigma-Aldrich (Johannesburg, South Africa), and CaO is provided by Associated Chemical Enterprises (Johannesburg, South Africa). The manganese ore is a pyrolusite (MnO₂)-based ore sourced from a small-scale ore body complex in South Africa, and the medium-volatile coal is also sourced from South Africa. The composition of the manganese ore and coal ash is shown in

Table 2. Bulk chemical composition of manganese ore and coal ash (mass%).

	%FeO	%MnO	%Cr2O3	%V2O5	%TiO2	%CaO	%K2O	%P2O5	%SiO2	%Al2O3	%MgO	%Na2O	%BaO
Ore	10.70	61.59	0.03	0.04	0.05	0.23	0.21	0.01	5.43	2.61	0.11	0.00	1.85
Ash	1.51	0.02	0.05	0.05	2.16	2.46	0.51	1.19	51.44	39.35	0.68	0.30	0.27

. The coal contains 22.5% volatile matter, 12% ash, 62.2% fixed carbon, and 3.3% moisture. The mixture, based on 100 g of ore, was placed in a graphite crucible and sealed with a graphite lid. The crucible dimensions were 50 mm inner diameter, 70 mm in height, and 7 mm in wall thickness. The lid was recessed by 7 mm to fit the crucible’s inner diameter. The reaction mixtures are summarised in

Table 3. Reaction mixtures (grams).

	Ore	Coal	Na2O.SiO2	CaO	Al	Si
Base case (BC)	100	27	20	15	0	0
A	100	10	20	15	30	0
C	100	10	20	15	30	10

.

2.2. Methods

The reaction mixture masses were weighed and mixed thoroughly in a mixing container. The mixture was transferred into the graphite crucible, and the graphite lid was placed onto the crucible rim. The muffle furnace was pre-heated and soaked for 12 h at the set-point temperature of 1350 °C. The furnace temperature was controlled using a proportional-integral-derivative (PID) controller. The aluminothermic reaction consists of placing the graphite crucible and contents into the muffle furnace. From prior experimentation, it was noted that the crucible must be removed from the furnace 2 min after ignition, marked by a flash flame exiting the muffle furnace’s top side panel joint gaps. Therefore, the reaction mixture was not entirely heated by the muffle furnace, but rather via the exothermic aluminium (aluminothermic) reaction. The ignition time was 6 min after placing the crucible in the furnace. Following a total reaction time of 8 min, the crucible and its contents were removed from the furnace. The lid was then removed, and the fluid metal and slag were poured into a cast-iron mould. Upon cooling to room temperature, the poured contents were returned to the graphite crucible for manual separation of metal and slag.

The large slag and alloy blobs were sectioned for chemical analysis of the alloy and slag phases using a scanning electron microscope (SEM). The cross-section faces were coated with gold for analysis by SEM via a Zeiss 540 Ultra FEG (field emission gun) SEM (Zeiss, Oberkochen, Germany) with energy dispersive X-ray spectroscopy (EDX) probe operated at 20 kV. Inductively coupled plasma optical emission spectroscopy (ICP-OES) method using a Perkin Elmer Optima 5300 instrument (Perkin Elmer, Springfield, IL, USA) was used in bulk alloy and slag analyses. The alloy carbon and sulphur content was analysed by the combustion method using a LECO CS 744 instrument (LECO, St Joseph, MO, USA).

3. Results

The base case sample did not form any significant metal within the reaction time, and at the low temperature of the muffle furnace (1350 °C), the crucible contents consisted of a loosely sintered mixture. The magnetic fraction of this sample is shown in Figure 1a. Figure 1b,c show that large alloy blobs were effectively separated from the slag in samples A and C, with some smaller metal droplets. It is seen that sample A formed one large alloy volume, and sample C formed multiple smaller alloy volumes. Thus, in sample A with added C and Al, and sample C with added C, Si, and Al, clear slag and alloy formation and separation were achieved.

EDX was applied to measure the alloy composition from the fractured surface made through the large alloy volume. A typical analysis area consisted of 0.5 mm × 0.5 mm, as shown for samples A and C alloy in Figure 2, Figure 3 and Figure 4; the analysis of these areas is summarised in

Table 4. EDS analyses of Sample A alloy (mass%); n.a. = not available; n.d. = not detected.

Area	%Al	%Si	%Mn	%Fe	%Ca	%V	%S	%C
1	0.6	2.4	69.4	26.9	0.1	0.6	0.0	n.a.
2	0.0	2.2	68.8	28.3	0.2	0.6	0.0	n.a.
3	1.6	2.3	68.2	27.0	0.2	0.6	0.0	n.a.
4	0.8	2.1	68.1	28.3	0.1	0.6	0.0	n.a.
5	1.4	1.8	67.5	28.6	0.2	0.6	0.0	n.a.
Average	0.9	2.2	68.4	27.8	0.2	0.6	0.0	n.a.
Maximum σ	0.01	0.01	0.06	0.03	0.00	0.01	n.d.
Bulk Alloy	0.4	1.4	66.1	28.0	0.05	0.40	0.03	3.5

and

Table 5. EDS analyses of Sample C alloy (mass%); n.a. = not available.

Area	%Al	%Si	%Mn	%Fe	%Ca	%S	%C
6	0.9	2.7	71.9	23.8	0.2	0.0	n.a.
7	1.1	3.0	71.4	23.7	0.4	0.0	n.a.
8	1.0	5.5	69.1	23.3	0.2	0.2	n.a.
9	1.1	6.8	67.9	23.0	0.2	0.3	n.a.
10	0.8	7.2	68.0	22.7	0.1	0.4	n.a.
Average	1.0	5.0	69.7	23.3	0.2	0.2	n.a.
Maximum σ	0.01	0.01	0.08	0.04	0.01	0.01
Bulk Alloy	1.4	8.7	65.8	21.8	0.007	0.08	2.2

. The bulk alloy analyses are also presented in Table 4 and Table 5, particularly the %C, as it is not accurately analysed by SEM-EDX. The analyses per area in each sample are closely matched. The main difference is evident in the increased silicon content in some regions of sample C (areas 8–10), most likely due to the incorporation of Si metal powder into the bulk alloy volume. The element maps in Figure 2 and Figure 3 indicate that the alloy matrix primarily consists of a Mn–Fe–Si alloy, with Al, Ca, and C appearing as specks. The C appears as surface flakes rather than being dissolved into the alloy. This is better seen in Figure 2b at the higher magnification of 2000X. The element maps for sample C in Figure 4b appear similar to those for sample A in Figure 2 and Figure 3. The slight difference is that no V was analysed. Some S was also analysed as displayed in Table 5 and confirmed to be present in both sample A and sample C bulk alloy analyses at 0.03% S and 0.08% S, respectively. The bulk alloy %C decreased with increased %Si, with 3.5% C at 1.4% Si in sample A alloy and 2.2% C at 8.7% Si in sample C alloy. The low Al content in both alloys confirms that the aluminium content addition in Table 3 is close to optimal. Too little Al addition would cause insufficient reduction of the ore, and too much Al addition would cause a high alloy %Al.

4. Discussion

Aluminothermic reduction is a non-equilibrium process due to the rapid release of heat, resulting in a short reaction time and limited room for process control, as observed in most furnace-based processes. As explained in prior work, the feed material’s recipe chemistry and particle size largely determine the final product outcomes, which can easily reach 3000 °C in the reaction melt [12,14,15,19]. Thus, calculating the feed recipe is not as straightforward as in most furnace-based processes.

4.1. Alloy Composition

From the %Al, %Si and %C analyses of sample A and sample C alloys, it is already clear that equilibrium is not a precise guide of the gas–slag–metal process in aluminothermic reduction. The carbon saturation level for sample A alloy at 1550 °C is 6.5 %C, and for sample C alloy with a higher %Si, the carbon saturation level is 4.2 %C. Compared to the %C shown in Table 4 and Table 5, at 3.5 %C in sample A alloy and 2.2 %C in sample C alloy.

The EDX results in Figure 2, Figure 3 and Figure 4 indicate that the alloys consist of a single phase with no apparent regular second phase present. The phase–chemical composition of each alloy was calculated using Thermo-Calc 2023b thermochemical software with the SGTE Alloy 4.9 database [20]. The bulk alloy compositions in Table 4 for sample A and Table 5 for sample C served as the basis for the calculations. As shown in Figure 5a, the sample A alloy composition can form multiple solid phases below the solidus temperature of 1093 °C. The liquidus temperature is close to the solidus temperature, at 1171 °C. The importance of a low liquidus temperature and a narrow solidification temperature range is that the alloy separation time from the melt is longer, allowing the alloy composition to be set in a short period. The phase chemistry for sample C, shown in Figure 5b, indicates fewer phases below the solidus temperature. Sample C has a lower solidus temperature of 975 °C and a liquidus temperature of 1099 °C, due to its higher silicon content of 8.7%, compared to only 1.4% Si in Sample A.

Although C was added in the form of coal, and ample C is available from contact with the graphite crucible, the take-up of carbon in the alloy is limited. The addition of coal as a reductant is intended to facilitate rapid pre-reduction to the MnO level, as previous studies have highlighted the need for pre-roasting MnO₂ ore to lower oxides and limit excessive heat release [11]. The synergistic effect of Si, acting both as a reductant and influencing alloy carbon uptake, is evident from the reduction work performed by Si and the expected decrease in carbon uptake in the Fe–Mn alloy with increased Si content [2]. This effect is shown in Figure 6a,b for the Mn-Si-C system at 1250 °C and 1450 °C. It is seen that for the phase boundary of liquid and graphite, the carbon saturation level of the liquid phase decreases with an increase in %Si. For example, in Figure 6a, the phase boundary of liquid and graphite coordinates varies from 5.2% C at 8.7% Si to 3.4% C at 13.4% Si. This effect is shown more clearly in Figure 6b when fewer solid phases are stable.

Recent work has shown that the contact time with the carbon crucible can be adjusted to vary the carbon content to some extent in steel production via aluminothermy, so preventing saturation of the steel with C at the reaction temperature [14]. Additionally, when using a non-carbon crucible in the carbothermic reduction of manganese ore, the alloy carbon content can be limited by controlling the feed mixture formulation and the slag-ore reaction layer [21,22].

4.2. Slag Composition

The target slag composition was selected to maintain a low liquidus temperature in the slag, even as a high level of Al₂O₃ uptake appears from the aluminothermic reduction reactions, such as reaction (1). Comparison of the average SEM slag analyses to the target slag analysis in

Table 6. Target vs. test slag analyses (mass%).

	%Al2O3	%SiO2	%MnO	%FeO	%CaO	%Na2O	%BaO	B2 = %CaO/%SiO2
Target slag	39	17	14	0.0	17	11	2	1.0
Slag A: EDX	54.5	9.1	20.0	0.0	8.0	5.5	2.8	0.9
Slag A: bulk	53.2	12.3	27.6	2.4	7.1	4.6	1.5	0.6
Slag A: Corrected	52.1	11.9	22.7	0.0	7.0	4.5	1.5	0.6
Slag C: EDX	52.4	18.7	6.5	0.1	12.6	7.5	2.3	0.7
Slag C: bulk	45.7	23.6	26.9	4.1	8.0	5.6	1.2	0.3
Slag C: Corrected	49.0	21.9	12.6	0.0	8.7	6.1	1.3	0.4

shows some differences from the expected composition. Comparison of the slag A composition to the target slag composition reveals lower percentages of CaO, Na₂O, and SiO₂, and higher levels of Al₂O₃ and MnO than expected. The differences may be due to some volatilisation of Na₂O, some reduction of SiO₂ and lesser reduction of MnO. Comparison of the slag C to the aim of slag composition also indicates lower percentages of Na₂O and CaO, as well as significantly lower percentages of MnO at 7%. This is partly due to the dilution of the slag with SiO₂ formed through silicothermic reduction using the added Si metal reductant.

3(MnO) + 2(Al) = <Al₂O₃> + 3(Mn)

(1)

< > = solid; ( ) = liquid

Despite the silicothermic reduction via Si, more Si was added to the alloy at 9% Si, as compared to the sample A alloy at 1% Si. Complex simultaneous reactions in the reaction system make pre-calculation of the alloy composition difficult, as the process is not an equilibrium process [14]. Therefore, the assumptions used in setting the target slag composition calculation are estimates, as is clear from the slag compositions in Table 6. The higher %Al₂O₃ content compared to the target slag %Al₂O₃ may indicate a preferential use of Al as a reductant, rather than C and Si.

In addition, the bulk chemical analyses of the product slags are displayed in Table 6. Reduction of iron oxides to metallic iron is expected to be complete. Therefore, the iron in the bulk analyses in Table 6 is considered to be due to alloy entrapment in the slag. The bulk slag analyses were corrected in terms of %MnO, %SiO₂, and %Al₂O₃ by subtracting the entrained alloy Mn, Si, and Al from the bulk slag analyses, and are labelled as “corrected” in Table 6. The total alloy yield consists of the weighed alloy mass and the slag-entrapped alloy mass, divided by the mass of Mn, Fe, and Si metal in the feed. The values are summarised in

Table 7. Mass balance numbers.

	Gram Weighed Alloy	Gram Mn in Weighed Alloy	Gram Fe in Weighed Alloy	Gram Weighed Slag	Gram Fe in Slag	Gram Alloy in Slag	Gram Mn in Ore	Gram Fe in Ore	Gram Mn + Fe + Si in Feed	%Alloy Yield
Sample A	18.7	12.4	5.3	79.1	1.5	5.3	47.7	8.3	56.0	43
Sample C	19.2	12.6	4.2	121.8	3.9	18.0	47.7	8.3	66.0	56

. It is observed that the yield of sample A alloy is 43%, compared to a yield of 56% for sample C alloy. These numbers are lower than the metal yield numbers of 58–64% reported for 10 kg of pre-roasted ore in the feed material mixture, as may be expected, since a larger volume of reacted material typically yields increased process efficiencies [11]. The element accounting percentages in

Table 8. Mass accounting of main elements (per cent).

	%Fe	%Mn	%Si	%Al
Sample A	81	67	58	62
Sample C	97	74	73	77

indicate a good accounting of Fe, with some loss of Mn, Si, and Al to the gas due to volatilisation of metals at the high reaction temperatures expected in thermic reduction. This effect may be countered by applying industrial methods, such as using a reaction chamber hood with a positive off-gas pressure to counter volatilisation and capture fumes for recycling [6].

The viscosities and phase chemistries of the slags were calculated using FactSage 7.3 thermochemical software, which utilised the FToxid database, as discussed in the following section. Based on the alloy yield values in Table 7, the addition of Si metal as a reductant aids overall alloy formation. However, a proportionally greater amount of alloy appears to be trapped in the slag. Figure 7 shows the calculated viscosity values of the slags at temperatures ranging from 1650 °C to 1900 °C. The slag liquidus temperatures are as follows: target slag at 1373 °C, slag A at 1652 °C, and slag C at 1530 °C. The higher SiO₂ containing slag in sample C is more viscous than the target slag and the sample A slag. This difference can account for the higher fraction of entrapped alloy in slag C. However, the viscosity values below 3 Poise should be all tappable slags in an industrial setting [23].

Since the precise target slag chemistry was not achieved in the product slag, the phase chemistry should be adjusted by increasing the Na₂O flux to transform the high Al₂O₃ content of the slag into the desired compound, NaAlO₂. The cooling curve of the target slag is shown in Figure 8a. It is seen that the slag liquidus temperature is 1373 °C and the slag solidus is much lower at 1010 °C, providing ample time for the alloy to separate from the slag upon cooling. It is observed that the desired leachable NaAlO₂ compound is expected to form upon cooling, although only up to a maximum of 37%. However, after tapping the slag from the aluminothermic reaction vessel, the slag can still be conditioned by adding more Na₂O to convert more Al₂O₃ to NaAlO₂. This effect is illustrated in Figure 8b for a mass of 20 kg of Na₂O.SiO₂ addition/100 kg slag, the proportion of NaAlO₂ is increased to 60% at the slag solidus temperature of 978 °C.

In conclusion, the results demonstrate that the sodium-oxide slag formulation used in the aluminothermic reduction of manganese ore offers an alternative, sustainable processing option for the circular economy. A lower %C containing ferro-manganese alloy can be produced in this way at less than the HCFeMn but higher than the MCFeMn carbon level. The addition of Al and Si in the alloy product may be beneficial, as the alloy can be considered a complex alloy for steelmaking alloying additions [5,6].

5. Conclusions

The alternative sustainable metallurgy processing route of sodium silicate fluxed aluminothermic reduction of manganese ore is illustrated with good slag–alloy separation. The Na₂O fluxed formulation enables the formation of a water-soluble NaAlO₂ compound, allowing for the recycling of Al₂O₃ through hydrometallurgy processing in the Bayer process. This process produces Al(OH)₃ feed for the Hall–Héroult electrochemical process, which is used to produce aluminium. In this way, CO₂ emissions are minimal if electricity input is sourced from non-fossil fuel energy sources. The addition of a small quantity of coal as a reductant facilitates rapid pre-reduction to the MnO level, thereby eliminating the need for an extra step of pre-roasting the MnO₂ ore. Analyses of the separated bulk alloy show compositions of 66% Mn, 22–28% Fe, 2–9% Si, 0.4–1.4% Al, and 2.2–3.5% C. This complex alloy can be used as a steelmaking alloy addition.