Enhancing the Selective Reduction of Nickel to Prepare FeNi50 Alloy from Saprolite-Type Laterite by CO-CO₂ Gas Pretreatment

Abstract

Owing to the superior reduction kinetics of limonite and goethite relative to silicates, coupled with the poor beneficiation performance of saprolite-type laterite, the direct carbothermal reduction of saprolite-type laterite exhibits limited nickel selectivity. This study leverages the selective oxidation effect of CO-CO₂ atmosphere on the metallic iron of pre-reduced minerals, as well as its suppression of Fe²⁺ reduction, to promote iron migration from oxides to the silicate phase, achieving homogenization and thereby negating its kinetic advantage in reduction. Parameter optimization experiments revealed that treating pre-reduced minerals with a 30 vol% CO atmosphere at 1200 °C for 20 min achieves complete iron homogenization within the silicate phase. Compared with the nickel–iron alloy (containing less than 10 wt% Ni) obtained via the RKEF process, the combination of pre-reduction, CO-CO₂ treatment, and the melting reduction process yielded nickel–iron alloys with nickel contents of 52.1 wt% (FeNi50 alloy) and 64.2 wt% at carbon consumptions of 4.0 wt% and 3.83 wt%, respectively, accompanied by nickel recovery rates of 95.5% and 91.2%. Furthermore, the enrichment of Fe²⁺ in the slag significantly reduces its melting point to approximately 1450 °C, enabling complete slag–metal separation after smelting at 1550 °C for 10 min.

1. Introduction

Nickel exhibits excellent chemical and physical properties, rendering it extensively utilized in diverse applications, including high-strength steel, stainless steel, high-temperature alloys, catalysts, and battery technologies [1,2,3,4,5,6]. As a high-demand element, global nickel production attained 3.7 million metric tons in 2024 [7]. Given the limited discovery of new nickel sulfide deposits in recent years, approximately 70% of current nickel production is extracted from laterite nickel ores, with this proportion demonstrating a consistent upward trend [8].

Laterite nickel ores represent the weathering profile of ultramafic protolith [9,10]. In tropical and subtropical regions with an annual precipitation exceeding 1000 mm, the dissolution of silicon, magnesium, and partial nickel from the protolith into groundwater occurs, while residual iron oxides accumulate on the surface, forming limonitic laterite nickel ores [11]. At the groundwater table interface, nickel combines with silicon and magnesium to generate hydrated silicates, accompanied by limonite and protolith residues, thereby constituting saprolite-type lateritic nickel ores [12,13]. Limonitic laterite nickel ores are predominantly utilized in hydrometallurgical processes like acid leaching and ammonia leaching or processed in selective reduction; saprolite-type laterite nickel ores are mainly processed by pyrometallurgical processes like RKEF and DRMS to produce iron–nickel alloy.

Recent research endeavors in hydrometallurgical processes have primarily concentrated on enhancing the leaching efficiency of nickel, elevating the recovery rate of nickel, and, most critically, mitigating environmental pollution. For instance, several studies have demonstrated the stress induced by techniques such as roasting or microwaving at a temperature of 600–800 °C for 1–2 h. Irradiation can effectively disrupt mineral structures, consequently enhancing nickel leaching efficiency by a factor of 2–4 [14,15,16,17]. The incorporation of surfactants of 2 × 10⁻⁵ mol/L of sodium dodecylbenzene sulfonate has been demonstrated to elevate the recovery of nickel from 85.2% to 99.6% [18,19], and mechanical agitation of 300 r/min [20,21,22] is wildly applied to mitigate nickel losses during the neutralization process by more than 90%. In the aspect of environmental protection, the organic acid leaching process was initially favored by researchers owing to its minimal environmental impact. Although recent research findings indicate that the leaching efficiency of organic acids like citric acid and oxalic acid falls short of meeting industrial production requirements [23,24,25], amidst the prevailing environmental protection initiatives, this technology demonstrates considerable potential for future development. Moreover, organic solvents have been proven to leverage the mass-to-charge ratio disparity of ions to selectively leach nickel to reduce acid consumption [26]. The alkaline leaching process has also demonstrated a similar leaching efficiency to that of the acid leaching process [14,27].

Conventional pyrometallurgical approaches like the rotary kiln electric furnace (RKEF) process, the use of a blast furnace, and direct reduction magnetic separation (DRMS) lack the selective nickel enrichment capability of hydrometallurgical processes [28,29]. However, they can provide metallic nickel to steel and alloy industries, which account for 70% of global nickel consumption, with lower cost. For instance, the rotary kiln electric furnace (RKEF) process, which accounts for approximately 80% of laterite nickel ore consumption, provides metallic nickel with 200$ lower per ton. Nevertheless, due to the physical beneficiation method demonstrating limited effectiveness in enhancing the Ni/Fe ratio of raw materials [30], the RKEF process is predominantly employed for smelting saprolite-type laterite nickel ores characterized by a nickel content exceeding 1.6 wt%, a SiO₂/MgO ratio ranging from 1.6 to 1.9, and an Fe/Ni ratio below 10. Since there are fewer and fewer high-nickel-grade minerals, the majority nickel–iron alloy products contain nickel ranging from 8 wt% to 12 wt% [31,32], which is insufficient to directly produce steel and alloys. The nickel content of nickel matte produced via the sulfide reduction roasting oxygen-enriched blowing process is comparable to that achieved through hydrometallurgical methods. However, the elevated production costs associated with the sulfidation stage, coupled with a relatively low nickel recovery rate (approximately 80%), render this approach economically comparable to hydrometallurgical methods [31].

Owing to the progressive depletion of high-grade lateritic nickel ores, the nickel concentration in alloys produced via conventional pyrometallurgical routes typically ranges from 8 to 12 wt%. To meet specifications for high-nickel stainless steels (e.g., >15 wt% Ni), these alloys must be blended with electrolytic nickel. However, producing electrolytic nickel via hydrometallurgical processing incurs a premium of approximately USD 2000–3000 per tonne of nickel contained relative to pyrometallurgical nickel production [31]. Consequently, elevating the nickel grade of primary pyrometallurgical alloys would substantially reduce the overall production cost of high-nickel ferroalloys and specialty steels. In order to enhance the selective reduction of nickel, some studies have applied sulfides or sulfates to convert FeO into FeS, thereby inhibiting the iron metallization process and elevating the nickel grade of the alloy by a factor of 1.5–3 [33,34,35,36]. Nevertheless, the associated costs of additives and nickel loss during the sulfidation process constrain this method’s industrial application. Hydrogen plasma can rapidly heat a mineral to its melting point within 5 min, which significantly mitigates the kinetic reduction advantage of oxides. However, as reported by U. Manzoor et al. [37], the selective reduction capability of hydrogen plasma towards nickel failed to attain the theoretically anticipated value necessary to yield high-grade ferronickel (70–90 wt% Ni content), with corresponding total recovery rates ranging from 60 % to 90%. After multiple reduction steps, the nickel content of the alloy decreased to 30 wt%, with a nickel recovery rate of 78%. It must be noticed that under non-molten reduction conditions, the reduction rate of limonite (the predominant iron-bearing phase in lateritic nickel ores) substantially exceeds that of silicate (the principal nickel-bearing phase in lateritic nickel ores) [28]. Given that the reduction process commences prior to the mineral’s melting phase, the actual reduction mechanism of laterite nickel ore inevitably deviates from the theoretically predicted homogeneous phase. As indicated by some research, the enhanced sintering process of iron with silicate induced by adding fluxes such as CaO, CaF₂, and NaCl deteriorates the iron reduction mechanism of oxide, thus enhancing the selective reduction of nickel [28,38,39]. Nevertheless, these approaches present an insufficient transformation of iron oxide, yielding nickel alloy concentrations of less than 20%.

Notably, the CO-CO₂ mixed gas system exhibits remarkable performance in restraining the reduction of Fe²⁺. Several studies have attempted to utilize CO-CO₂ mixed gases to either inhibit the reduction of Fe²⁺ or to selectively oxidize the metallic iron generated through reduction processes. These approaches have demonstrated efficacy in suppressing the metallization of iron [40,41]. However, the reduction kinetics of metal ions in non-molten silicate systems remain relatively slow, and selective nickel reduction at lower temperatures through controlled atmospheric conditions may require prolonged processing duration. In contrast to previous investigations, the present study reveals that mixed gases facilitate the rapid conversion of iron oxides into iron silicate under non-molten conditions, thereby elevating the kinetic barriers for iron reduction. Owing to the substantially lower Gibbs free energy required for nickel reduction compared to ferrous ions, this study has demonstrated that nickel can be rapidly and selectively reduced to metallic nickel during the molten reduction stage of mixed gas-treated minerals.

2. Materials and Methods

2.1. Material

The raw material of this research involves saprolite-type lateritic nickel ore from Sulawesi Island, Indonesia, and anthracite from Shanxi China.

Table 1. Composition of saprolite-type laterite nickel ore (wt%).

Fe3+	Fe2+	T.Ni	SiO2	MgO	Al2O3	CaO	T.Cr	T.Mn	T.Co	IL
15.89	0.56	1.46	42.06	15.23	2.73	0.79	0.88	0.35	0.045	14.33

presents the chemical analysis of the dried materials (desiccated at 120 °C until reaching constant weight). The quantification of non-oxygen elements was conducted through ICP-AES and XRF, while Fe²⁺ content was ascertained via the titration method. The mineral represents a low-grade saprolite-type lateritic nickel ore, characterized by a SiO₂/MgO molar ratio of 2.9 and a Fe/Ni mass ratio of 11.3 (Figure 1a). The industrial production data pertaining to this mineral demonstrates that without refining treatment, the RKEF process could only yield nickel–iron alloys with a nickel content below 10 wt%. The desiccated anthracite exhibited a fixed carbon content of 69.92 wt% (

Table 2. Occurrence states of Ni and Fe in saprolite-type laterite nickel ore (%).

Distribution	Sulfides	Oxides	Silicates	Manganese
Ni	0.4	19.3	79.3	1.0
Fe	0.8	61.1	37.7	0.4

). The occurrence states of iron and nickel in the mineral were determined through chemical phase analysis [42], with the corresponding results presented in

Table 3. Composition of anthracite (wt%).

C	Ca	Mg	Si	Al	Fe	Volatiles
69.92	1.75	0.79	5.71	0.97	1.82	8.75

.

The analytical data indicates that nickel and iron predominantly exist in oxide and silicate forms, respectively, wherein approximately 80% of nickel is present in silicates and about 60% of iron is found in oxides. The phase analysis and elemental distribution of the mineral were characterized by XRD (Bruker, Karlsruher, Germany) and SEM-EDS (Oxford Instruments, Oxford, UK), as illustrated in Figure 1a,c. Mineralogical analysis revealed that the specimen predominantly consists of quartz, limonite, and silicate phases. Due to the isomorphic occurrence state of nickel with iron within oxides and iron and magnesium within silicates [28,40], no discrete nickel-rich mineral phases was observed. Given the complex phase assemblage of the specimen and the intensive picks of all characteristic XRD peaks concentrated within the 10° to 60° 2θ range, the XRD analysis presented in this study is confined to this angular range to facilitate clearer interpretation and comparison. The TG-DSC curves of the minerals are presented in Figure 1b. Owing to capillary action, the adsorbed water is released at 113 °C. The dehydration of goethite occurs at 315 °C, while that of silicate takes place at approximately 636 °C. The endothermic peak observed at 818 °C corresponds to the recrystallization process of certain silicate minerals.

2.2. Method

Specimen preparation procedure: The dried ore and anthracite were pulverized to a particle size of less than 18 mesh and homogenously blended at a mass ratio of 100:9.5. Subsequently, the mixture was compacted into cylindrical specimens with a diameter of 20 mm at a pressure of 10 MPa. The prepared specimens exhibited a carbon content of 6.1 wt%, with a molar ratio of fixed carbon to oxygen in iron oxides and nickel oxide of 1.2:1.

Pre-reduction oxidation procedure: The specimen (300 g ± 0.1 g per batch) is positioned within a nickel–chromium alloy basket and subjected to thermal treatment in an alumina tube furnace (Wuhan Electric Furnace Instrument, Wuhan, China), which is subsequently heated at a rate of 30 °C/min until reaching the designated reduction temperature. Throughout the reduction phase, ultra-high-purity nitrogen gas (99.999%) is introduced into the tube furnace. Upon completion of the pre-reduction, the tube furnace temperature is adjusted to the oxidation temperature at a rate of 50 °C/min, during which the ultra-high-purity nitrogen gas is replaced with an oxidizing gas mixture for isothermal oxidation. Following oxidation, the oxidizing gas is switched back to ultra-high-purity nitrogen gas, and the specimen is furnace-cooled to ambient temperature prior to removal.

Smelting procedure: The smelting specimens are positioned within an alumina crucible and subjected to smelting in a carbon tube furnace. The furnace is heated at a controlled rate of 100 °C/min until reaching 1600 °C, maintaining this temperature for a duration of 10 min. Subsequently, the specimen is furnace-cooled to ambient temperature, after which they are extracted for subsequent slag–metal separation. Throughout the entire process, a continuous flow of high-purity nitrogen is maintained.

For each specimen, four independent component analyses were performed, and the average value was adopted as the analytical result. Specifically, the determination of metallic iron and nickel in the pre-reduction and oxidized specimens was conducted through the alkali eutectic and ferric chloride leaching methodology [43]. The carbon content in the sample was determined by a carbon–sulfur analyzer. The metallization degree of nickel (M_Ni) and iron (M_Fe) of pre-reduced specimens and oxidized specimens are determined by Equations (1) and (2), respectively:

$${\mathit{M}}_{\mathrm{N}\mathrm{i}}=({\mathrm{w}}_{\mathrm{s}}({\mathit{M}}_{\mathrm{N}\mathrm{i}})/{\mathrm{w}}_{\mathrm{s}}({\mathit{T}}_{\mathrm{N}\mathrm{i}}))\times 100\%$$

(1)

$${\mathit{M}}_{\mathrm{F}\mathrm{e}}=({\mathrm{w}}_{\mathrm{s}}({\mathit{M}}_{\mathrm{F}\mathrm{e}})/{ \mathrm{w}}_{\mathrm{s}}({\mathit{T}}_{\mathrm{F}\mathrm{e}}))\times 100\%$$

(2)

where w_s(T_Ni) and w_s(T_Fe) denote the mass fractions of total nickel and total iron within specimens, while w_s(M_Ni) and w_s(M_Fe) represent the mass fractions of metallic nickel and metallic iron within specimens, respectively. The oxidation degree of nickel (O_Ni), iron (O_Fe), and carbon (O_C) are determined by Equations (3)–(5), respectively:

$${\mathit{O}}_{\mathrm{N}\mathrm{i}}=(({\mathit{M}}_{\mathrm{N}\mathrm{i}}\mathit{R}-{\mathit{M}}_{\mathrm{N}\mathrm{i}}\mathit{O})/{\mathit{M}}_{\mathrm{N}\mathrm{i}}\mathit{R})\times 100\%$$

(3)

$${\mathit{O}}_{\mathrm{F}\mathrm{e}}=(({\mathit{M}}_{\mathrm{F}\mathrm{e}}\mathit{R}-{\mathit{M}}_{\mathrm{F}\mathrm{e}}\mathit{O})/{\mathit{M}}_{\mathrm{F}\mathrm{e}}\mathit{R})\times 100\%$$

(4)

$${\mathit{O}}_{\mathrm{C}}=(({\displaystyle \frac{{\mathrm{w}}_{\mathrm{R}}(\mathrm{C})}{{\mathrm{w}}_{\mathrm{R}}(\mathrm{Mg})}-}{\displaystyle \frac{{\mathrm{w}}_{\mathrm{O}}(\mathrm{C})}{{\mathrm{w}}_{\mathrm{O}}(\mathrm{Mg})}})/{\displaystyle \frac{{\mathrm{w}}_{\mathrm{R}}(\mathrm{C})}{{\mathrm{w}}_{\mathrm{R}}(\mathrm{Mg})}})\times 100\%$$

(5)

where M_NiO and M_FeO refer to the metallization degree of nickel and iron within oxidized specimens; M_NiR and M_FeR denote the metallization degree of nickel and iron within pre-reduced specimens; w_R(C) and w_O(C) represent the mass fractions of carbon in pre-reduced specimens and oxidized specimens; and w_R(Mg) and w_O(Mg) represent the mass fractions of magnesium in pre-reduced specimens and oxidized specimens, respectively. The recovery rate of nickel (R_Ni) and iron (R_Fe) are calculated using Equations (6) and (7), respectively:

$$R_{\mathrm{Ni}}=({\mathrm{w}}_{\mathrm{a}}(\mathrm{Ni})\times {\mathrm{m}}_{\mathrm{a}})/({\mathrm{w}}_{\mathrm{sl}}(\mathrm{Ni})\times {\mathrm{m}}_{\mathrm{sl}}+{\mathrm{w}}_{\mathrm{a}}(\mathrm{Ni})\times {\mathrm{m}}_{\mathrm{a}})\times 100\%$$

(6)

$$R_{\mathrm{Fe}}=({\mathrm{w}}_{\mathrm{a}}(\mathrm{Fe})\times {\mathrm{m}}_{\mathrm{a}})/({\mathrm{w}}_{\mathrm{sl}}(\mathrm{Fe})\times {\mathrm{m}}_{\mathrm{sl}}+{\mathrm{w}}_{\mathrm{a}}(\mathrm{Fe})\times {\mathrm{m}}_{\mathrm{a}})\times 100\%$$

(7)

where w_sl(Ni) and w_sl(Fe) represent the mass fractions of nickel and iron for slag; w_a(Ni) and w_a(Fe) refer to the mass fractions of nickel and iron for alloy; and m_a and m_sl denote the mass of alloy and slag, respectively.

Other key parameters in this study encompass the relative nickel content in the alloy, denoted as C(Ni) (expressed as the mass percentage of Ni/(Ni + Fe)), the relative iron content in the alloy, designated as C(Fe) (represented as the mass percentage of Fe/(Ni + Fe)), the volumetric concentration of CO in the oxidizing atmosphere, φ(CO), and the volumetric concentration of CO₂ in the oxidizing atmosphere, φ(CO₂).

3. Results and Discussion

3.1. Carbothermic Reduction

Before elaborating on the CO-CO₂ treatment, it is essential to observe the carbothermal reduction of this mineral, particularly the kinetic constraints that impede nickel reduction. Minerals subjected to thermal treatment at 1000 °C are designated as the control group for comparative analysis. As depicted in Figure 2b, all goethite undergoes complete transformation into limonite, while most silicate hydrate transitions to an amorphous state due to the removal of crystalline water. Furthermore, given that the iron and nickel present in sulfides and manganese ores constitute approximately 1% of the total iron and nickel content, and considering the challenges in tracking the reduction process of these minerals, along with their negligible impact on the overall reduction process of iron and nickel, this study primarily focuses on the reduction process of iron and nickel contained within oxides and silicates. To facilitate the differentiation of iron and nickel in various states of existence, the iron and nickel in oxides are designated as Fe(O) and Ni(O), respectively, while those in silicates are denoted as Fe(Si) and Ni(Si) in subsequent discussions.

The inherited physical difference between oxides and silicates lead to significant discrepancies in their reduction rates. The oxides exhibit a loose and porous structure, with their porosity further enhanced following dehydration. The substantial specific surface area facilitates comprehensive adsorption and diffusion of the reducing agent on its surfaces, thereby enabling their rapid reduction [44,45]. The nickel–iron alloy formed during reduction acts as an effective catalyst for the Boudouard reaction, thereby increasing the concentration of active carbon and enhancing the overall reducing potential of the atmosphere to accelerate the reduction of oxides. Silicates not only exhibit a dense crystalline structure but also demonstrate a propensity for facile solid-state reactions with Fe²⁺ generated during the reduction process, which further impairs the reduction kinetics [41]. Given that approximately 20% of Ni and 60% of Fe in the specimens are present in oxide forms, it is reasonable to observe a significant decline in the metallization rate of Ni and Fe once Ni(O) and Fe(O) complete their reduction, as evidenced by the inflection points in the M_Ni and M_Fe curves when M_Ni and M_Fe reach 20% and 60%, respectively (Figure 2a).

As illustrated in Figure 2b, the metallization of Ni and Fe in the initial stage of the reduction process is predominantly governed by the reduction of oxides: the predominant proportion of oxides has been reduced within 10 min at 1000 °C. Within this stage, the transformation of Fe(Si) into Fe²⁺ catalyzed the crystallization of amorphous silicates, resulting in the formation of enstatite and clinoenstatite. Concurrently, Ni(O) is directly reduced to metallic nickel, while the predominant portion of Fe(O) is reduced to Fe²⁺ and subsequently adsorbed into silicates to form an iron-rich olivine crust enveloping the silicate matrix (Figure 3a). This phenomenon elucidates the significantly impeded reduction kinetics of Fe(O), resulting in the metallization process of Fe(O) lagging behind the reduction process of oxides. Consequently, within the specimen reduced at 1000 °C for 10 min, M_Ni attains 21.5%, whereas M_Fe reaches only 6.9% (Figure 2a). The alloy product primarily consists of tetrataenite (with a nickel content ranging from 27 wt% to 65 wt%) (Figure 2b).

Nevertheless, the olivine crust formed in the initial stage of the reduction exhibits a markedly larger specific surface area than the silicate matrix and preferentially initiates contact with the reducing agent, thereby presenting substantially superior reduction kinetics. As depicted in Figure 2a, following the complete reduction of oxides, the metallization rate of nickel undergoes a notable deceleration (the inflection point of M_Ni curve), whereas iron maintains a relatively high metallization rate. As the reduction proceeds, both the temperature and duration increase, and the metallization rate of iron significantly increases, subsequently exceeding and diverging from that of nickel until the complete reduction of olivine, where the inflection of M_Fe curve is observed. From a microscopic perspective and the XRD analysis, as depicted in Figure 2b and Figure 3b, the olivine experiences a markedly faster reduction than the silicate matrix; as the reduction temperatures increase, olivine progressively transfers into metallic iron (wherein tetrataenite assimilates metallic iron and transforms into an iron–nickel alloy), cristobalite, and enstatite, while little change in clinoenstatite content is observed. In the specimen reduced at 1200 °C for 30 min, olivine completes its reduction, leaving a metal–silicate transition layer with an approximate thickness of 3 μm (Figure 3a). Concurrently, M_Fe and M_Ni achieve 64.2% and 40%, respectively.

It is evident that the differential occurrence states of iron and nickel in mineral structures impose significant constraints on the reduction kinetics of nickel in non-molten reduction. Although mineral homogenization occurs post-melting, achieving uniform dispersion of the reductant within the molten mineral prior to triggering the reduction process presents significant challenges. Direct introduction of the reducing agent onto the melt surface will confine the reduction reaction to the interfacial region where the melt and reducing agent interact. When the reduction rate of Ni²⁺ surrounding the reductant exceeds the rate of Ni²⁺ diffusing from the melting into the interfacial zone, the reduction rate of nickel decreases. Following the application of 4 wt% carbon powder onto the molten mineral surface at 1600 °C with a 10 min holding period, the recovery rate of Ni (R_Ni) and Fe (R_Fe) were measured at 56.5% and 35.1%, respectively, indicating that the reduction advantage of nickel remains unremarkable.

In conclusion, at least two fundamental prerequisites are essential for achieving selective nickel reduction in minerals: deep homogenization of iron within the mineral matrix and relatively uniform distribution of the reducing agent throughout the mineral. Chemical phase analysis revealed that following pre-reduction and subsequent selective oxidation, the Fe(Si) phase content increased from 37.7% to a maximum of 99.4%, which was achieved in the sample subjected to oxidation in a pure CO₂ atmosphere at 1200 °C.

3.2. CO-CO₂ Treatment

The oxidative potential of a CO₂-CO mixed gas exhibits significant variations in its interaction with metallic iron and metallic nickel. As depicted in Figure 4a (calculated using Factsage 8.2; Database: FactPs; input parameters: molar ratio; temperature range: 300–1300 °C, 1 atm), within the temperature range of 700 °C to 1300 °C, the oxidation resistance hierarchy follows the sequence: Ni > Fe₃O₄ > FeO > Fe > C. In an oxidation temperature of 1200 °C, to achieve complete inhibition of metallic iron oxidation, the volume fraction of CO (φ(CO)) must exceed 77.8%, whereas for complete inhibition of metallic nickel oxidation, a φ(CO) value above 2.4% is sufficient. This computation result shows an alignment with the experimental findings reported by Hang et al.: a gas mixture containing 70% CO (φ(CO) = 70%) failed to reduce Fe²⁺ in saprolite-type laterite nickel ore to metallic iron at a temperature of 1100 °C [38]. Consequently, from a thermochemical perspective, the application of this gas mixture can facilitate Fe²⁺ diffusion into the silicate matrix interior, thereby enhancing iron homogenization within the silicate matrix.

A pure CO₂ with a controlled flow rate of 2 L/min was employed to oxidize the pre-reduced specimens (which were pre-reduced at 1200 °C for 30 min) to observe the behavior of iron, nickel, and carbon in CO₂ atmosphere. While thermochemical analysis indicates that carbon exhibits greater susceptibility to oxidation by CO₂ compared to metallic iron, experimental data reveal that metallic iron demonstrates a higher oxidation rate than carbon (Figure 5). This discrepancy can be attributed to the predominant existence of fixed carbon located within the benzene ring structure of the reducing agent, where the dissociation of C-O-O-C bonds formed through CO₂ adsorption necessitates substantial activation energy [46]. As reported in relevant studies, the gasification reaction of carbon in conventional coal typically initiates at approximately 900 °C, with the gasification rate doubling for every 30 °C to 40 °C increase in temperature [47,48]. Consequently, as depicted in Figure 5a, elevating the oxidation temperature from 900 °C to 1200 °C, the O_Fe increases moderately from 88.77% to 95.96%, while the O_C surges from 11.65% to 85.42%.

In a weakly oxidative atmospheric environment, Fe²⁺ ions, generated through both reduction processes and the oxidation of metallic iron, are expected to continuously diffuse into the silicate matrix without undergoing reduction to metallic iron. The oxidation temperature influences both the stability of Fe²⁺, as illustrated in Figure 4a, and its diffusion into silicate. As demonstrated in Figure 6a and Figure 7a, at an oxidation temperature of 900 °C, the relatively weak stability of Fe²⁺, coupled with its sluggish diffusion process into the silicate matrix, renders Fe²⁺ susceptible to further oxidation. The majority of metallic iron within this oxidized specimen is oxidized into magnetite, thereby forming an iron-rich magnetite–olivine crust covering the silicates. As the oxidation temperature increased, both the stabilization of Fe²⁺ and their diffusion process enhanced, which consequently inhibited magnetite formation and promotes the formation of iron silicate. From the perspective of the XRD spectrum (Figure 6a), the magnetite phase is entirely absent at an oxidation temperature of 1100 °C, while the accelerated diffusion kinetics of Fe²⁺ facilitates the homogenization of iron into the silicate matrix, consequently forming silicates of olivine, pigeonite, and enstatite phases. Moreover, microscopic observation indicates that specimens oxidized at 1200 °C exhibit significantly improved sintering characteristics relative to those treated at 900 °C (Figure 8) and exhibit deep homogenization of iron within the silicate matrix.

Ni²⁺ is more susceptible to the reduction compared to Fe²⁺. Experimental results indicate (Figure 4) that a residual carbon content exceeding 0.5 wt% is adequate to maintain a reduction rate of Ni²⁺ that surpasses the oxidation rate of metallic nickel, consequently leading to a progressive increase in the M_Ni of the specimen. The marked disparity in oxidation resistance between metallic iron and metallic nickel substantially enhances the nickel enrichment within the alloy during the oxidation process. As demonstrated in Figure 7b and Figure 8a, in the specimen oxidized at 900 °C for 30 min, nickel displays a nearly identical concentration level to that of iron within the alloy matrix, while exhibiting a significantly higher concentration degree within the mineral phase compared to iron. With the elevation of oxidation temperature, the nickel content in the alloy demonstrates progressive enhancement; nevertheless, as the carbon component approaches complete oxidation, the metallic nickel inevitably experiences concurrent oxidation in pure CO₂ environment. Subsequent to a 20 min oxidation process at 1200 °C, the M_Ni of the specimen increased from 40% to 58.5%, whereas the carbon content decreased from 2.99 wt% to 0.54 wt%. Further decreased carbon content becomes insufficient to inhibit the oxidation of metallic nickel. Upon extending the oxidation duration to 30 min, the M_Ni decreased to 25.2%. Concurrently, the oxidation products of metallic nickel diffused into the silicate matrix, exhibiting a similar behavior to that of metallic iron oxidation products (Figure 7b).

From a structural perspective, the iron–nickel alloy exhibits a body-centered cubic lattice configuration, wherein the relatively larger interatomic spacing facilitates oxygen diffusion into the alloy matrix and hinders the tight bonding between oxide films and metal substrates. The oxide film primarily composed of iron oxide exhibits higher propensity for solid-state reactions with the silicate matrix compared to nickel oxide. As depicted in Figure 6, iron–nickel alloy presents preferential oxidation to transform into tetrataenite (tetragonal crystal structure) with increased nickel content and denser crystalline arrangement. Following the complete transformation of iron–nickel alloys into tetrataenite, the oxidation rate of iron undergoes a substantial reduction, subsequently leading to a progressive transformation of tetrataenite into a face-centered cubic nickel–iron alloy.

To investigate the effect of CO in the oxidation process, a mixed gas atmosphere consisting of CO₂ and CO with a controlled flow rate of 10 L/min was introduced. The entire oxidation process was maintained at 1200 °C for a duration of 30 min. As depicted in Figure 5c, the increased gas flow rate markedly accelerated the oxidation rate of the specimens: nearly all metallic nickel, metallic iron, and carbon underwent oxidation in an atmosphere with φ(CO) of 0%. Although CO demonstrates remarkable efficacy in inhibiting the oxidation of metallic nickel, the solid-state interaction between nickel oxidation products and the silicate matrix compromises its reduction kinetics. The experimental findings indicate that the φ(CO) necessary to achieve complete inhibition of metallic nickel oxidation substantially exceeds the thermochemical prediction. In specimens subjected to oxidation in a mixed gas with φ(CO) of 5%, a substantial quantity of fine alloy particles is observed, accompanied by a reduction in the O_Ni value to 54.3%. However, the residual alloy content remains insufficient to generate distinct diffraction peaks in the XRD spectrum (Figure 6b). Elevating the φ(CO) in the gas mixture progressively impairs its oxidative capacity, thereby inhibiting the oxidation of metal, with a corresponding reduction in the nickel content of the alloy (Figure 9a,b). As depicted in Figure 9a, specimen subjected to oxidation under an atmosphere with φ(CO) of 30% presents a substantially enhanced nickel concentration; however, is also presents reduced nickel content in the alloy compared to that oxidized at φ(CO) of 5%.

Although O_Ni transitioned to negative values when φ(CO) was increased to 20%, complete inhibition of nickel oxidation required φ(CO) to reach approximately 30%. This conclusion can be drawn from two observations: first, the O_Ni curve (Figure 5b) exhibits a continuous decline when φ(CO) reaches 20% but stabilizes when φ(CO) approaches 30%; second, upon extending the oxidation duration to 60 min, the O_Ni in the specimen oxidized at O_Ni(CO) of 20% decreased to 17.3%, whereas the specimen oxidized at O_Ni(CO) of 30% shows little deviation from its initial value. The negative O_Ni value observed in the specimen oxidized at O_Ni(CO) of 20% might result from the continuous reduction of Ni²⁺ by residual carbon during the initial oxidation stage, combined with the significantly reduced oxidation rate for metallic nickel in this atmosphere, thereby producing more metallic nickel than was oxidized during the process.

3.3. Selective Nickel Reduction

In silicates, oxygen exists predominantly as bridging oxygen (BO) and non-bridging oxygen (NBO), both of which exhibit low degrees of ionic dissociation during melting. Furthermore, the extensive polymerization of [SiO₄] tetrahedra in the silicate melt severely restricts the mobility and accessibility of Fe²⁺ and Ni²⁺ ions (hereafter denoted as Fe(Si) and Ni(Si), respectively), thereby suppressing their reduction reactivity. Consequently, direct melting of this mineral fails to yield a homogeneous melt. In contrast to the silicate phase, limonite undergoes extensive thermal decomposition during melting, characterized by dehydroxylation and collapse of its poorly ordered oxyhydroxide structure, yielding a melt with markedly higher ionic dissociation. Consequently, in the mixed melt comprising both phases, Fe²⁺ exhibits significantly greater reaction activity than Ni²⁺. Following the pre-reduction and selective oxidation treatment, Fe(O) homogenizes into the silicate matrix, thereby facilitating the formation of homogeneous melt. During smelting, transition metal carbides dissolved in the alloy undergo thermal decomposition, releasing reactive carbon species that reduce metal oxides in the slag. In a homogeneous melt, Ni²⁺ is preferentially reduced over Fe²⁺ due to its more negative standard Gibbs free energy change for reduction, consistent with thermodynamic predictions, leading to selective nickel extraction and a consequent enrichment of Fe²⁺ in the residual slag.

Given that the primary constituents of the slag generated during the smelting of saprolite-type laterite nickel ore predominantly comprise the FeO-MgO-SiO₂, its melting point characteristics can be roughly interpreted in Figure 4b. Despite the SiO₂/MgO ratio of the mineral reaching 2.9, the substantial increase in Fe²⁺ content in the silicate matrix during the iron homogenized process significantly reduces the melting point of the slag from about 1800 °C to below 1500 °C. Evidently, the significantly reduced melting point and elevated basicity of slag not only reduced smelting temperature and flux consumption but also facilitated the metal–slag separation. It should be noted that although the sample was melted in an alumina crucible, crucible erosion was negligible under the experimental conditions: the slag basicity was low (0.37–0.8), and the melting duration was brief. This is evidenced by the minimal incorporation of alumina into the slag, with the aluminum content increasing by less than 1 wt%. The experimental results, as illustrated in Figure 10a, demonstrate that during the smelting process of oxidized specimens, the metallization rate of nickel markedly exceeds that of iron: while maintaining R_Ni above 95%, R_Fe can be less than 10%. Among the smelting products, the highest relative nickel content (C(Ni)) in the alloy reaches 85.7%, albeit with an R_Ni of only 30.8%. To sustain an R_Ni of 95%, specimens oxidized at a φ(CO) of 30% produce alloys with a C(Ni) up to 48.2%. Additionally, all oxidized specimens were observed to completely melt below 1500 °C.

Nevertheless, from both economic and environmental perspectives, excessive reductant usage and over-reduction of minerals would lead to unnecessary energy consumption and increased carbon emissions. From an industrial production standpoint, the selective oxidation process presents challenges in precisely controlling the residual carbon content within the specimen, consequently impeding the accurate regulation of the nickel grade in the alloy. Given that the objective of the pre-reduction and selective oxidation process is to homogenize iron into the silicate phase, the following optimizations were implemented: reducing the carbon of the briquette from 6.1 wt% to 3.5 wt% (equal to 5 wt% anthracite), decreasing the reduction duration to 10 min, utilizing gas with φ(CO) of 30% as the oxidation gas, reducing the oxidation duration to 20 min, crushing the oxidized specimen and mixing it with the reducing agent prior to smelting, and lowering the smelting temperature to 1550 °C. The schematic representation of this process and the corresponding results are illustrated in Figure 10b. As anticipated, Figure 10b shows that the addition of carbon to the oxidized specimens during smelting did not change the Fe²⁺ in the silicate matrix to Fe; however, the selective reduction efficacy of nickel maintained exceptional performance following the process optimization: with the incremental addition of carbon from 0 to 0.2 wt% (relative to oxidized mineral), R_Ni increases from 30.8% to 70.8%, whereas R_Fe merely increases from 0.5% to 2.7%. Notably, the metallization rate of nickel is 18-fold higher than that of iron, while C(Ni) in the alloy attains 74.9 wt%. However, the reduction in Ni²⁺ concentration within the slag consequently leads to a decrease in the nickel reduction rate. When the carbon addition reaches 0.4 wt% and 0.6 wt%, R_Ni achieves 91.2% and 95.5%, while the C(Ni) of the alloy decreases to 64.2 wt% and 52.1 wt%, respectively. Further augmentation of carbon addition exerts a negligible impact on the enhancement of R_Ni, accompanied by a continuous decrease in the nickel grade of the alloy. The carbon balance for the 0.6 wt% carbon addition sample is shown in

Table 4. Carbon balance for sample with 0.6 wt% carbon addition.

(Per 100 g of Raw Ore)	Raw Briquette	Reduction	Oxidation	Carbon Addition	Melting
Carbon input	3.5	0	0	0.5	0
Carbon output	0	2.86	0.54	0	0.59
Residual carbon	3.5	0.64	0.1	0.6	0.01

.

It is important to highlight that the oxidized specimens must undergo carbon blending prior to the molten process. As previously mentioned, if the oxidized specimens are subjected to melting initially, followed by carbon spraying onto the slag surface, carbon would concentrate on the slag surface, thereby generating a Ni²⁺ depletion zone surrounding the reductant. Under an identical carbon addition of 1.0 wt%, the R_Ni of the prior carbon blending attains 97.9%, with the C(Ni) in the alloy reaching 33.4%. In contrast, the R_Ni and C(Ni) of the alloy in the latter case decrease to 69.4% and 21.8%, respectively.

Although Hang et al. [40] employed a comparable strategy to produce high-grade nickel–iron alloy from saprolitic laterite nickel ores, their investigation did not address carbon behavior or iron homogenization during pre-reduction and selective oxidation. In contrast, this study introduced an intensified nickel reduction step during pre-reduction, achieved by incorporating CaF₂, which promoted the extensive formation of metallic nickel in the mineral phase following selective oxidation. This process yielded two critical outcomes: first, it elevated the carbon solubility in the alloy; second, residual carbon exhibited substantially greater reducing capacity than Ni²⁺ in the slag during melting, thereby driving significant internal reduction of Fe²⁺. As a result, deep oxidation exerted only a marginal influence on further nickel enrichment in the final alloy.

4. Conclusions

The pre-reduction and CO-CO₂ treatment facilitates the migration of iron elements from their original oxide state into silicate structures, thereby achieving compositional deep homogenization. This process exerts a dual effect: first, it compromises the reduction kinetics of the iron within oxides, thereby enhancing the selective reduction of nickel; second, it reduces the eutectic point of the mineral matrix to below 1500 °C and elevates the basicity, which is critical for a thorough slag–metal separation, from 0.36 to above 0.8. The combination of the pre-reduction process, CO-CO₂ treatment, and secondary carbon addition smelting process has shown itself to be capable of producing a FeNi50 alloy with a nickel content of 52.1 wt% from low-grade laterite nickel ore (characterized by a nickel content of 1.46 wt%, a Fe/Ni ratio of 11.3, and a SiO₂/MgO ratio of 2.9) at a smelting temperature of 1550 °C, with a nickel recovery rate of 95.5% and total carbon consumption of 4 wt%. And, if the secondary carbon addition is reduced to 0.4 wt% (total carbon consumption of 3.83 wt%), the alloy nickel content increases to 64.2 wt%, while the nickel recovery rate decreases to 91.2%. This method has the potential to enhance raw material adaptability (expanded Fe/Ni ratio and SiO₂/MgO ratio) and selective nickel reduction (6.2-fold nickel concentration of alloy product relative to raw material) in the pyrometallurgical process.