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Article

Fe–P Alloy Production from High-Phosphorus Oolitic Iron Ore via Efficient Pre-Reduction and Smelting Separation

1
School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, China
2
Vale-CSU Joint Laboratory for Low-Carbon and Hydrogen Metallurgy, Central South University, Changsha 410083, China
*
Authors to whom correspondence should be addressed.
Minerals 2025, 15(8), 778; https://doi.org/10.3390/min15080778
Submission received: 18 June 2025 / Revised: 14 July 2025 / Accepted: 24 July 2025 / Published: 24 July 2025
(This article belongs to the Section Mineral Processing and Extractive Metallurgy)

Abstract

Diverging from conventional dephosphorization approaches, this study employs a novel pre-reduction and smelting separation (PR-SS) to efficiently co-recover iron and phosphorus from high-phosphorus oolitic iron ore, directly yielding Fe–P alloy, and the Fe–P alloy shows potential as feedstock for high-phosphorus weathering steel or wear-resistant cast iron, indicating promising application prospects. Using oolitic magnetite concentrate (52.06% Fe, 0.37% P) as feedstock, optimized conditions including pre-reduction at 1050 °C for 2 h with C/Fe mass ratio of 2, followed by smelting separation at 1550 °C for 20 min with 5% coke, produced a metallic phase containing 99.24% Fe and 0.73% P. Iron and phosphorus recoveries reached 99.73% and 99.15%, respectively. EPMA microanalysis confirmed spatial correlation between iron and phosphorus in the metallic phase, with undetectable phosphorus signals in vitreous slag. This evidence suggests preferential phosphorus enrichment through interfacial mass transfer along the pathway of the slag phase to the metal interface and finally the iron matrix, forming homogeneous Fe–P solid solutions. The phosphorus migration mechanism involves sequential stages: apatite lattice decomposition liberates reactive P2O5 under SiO2/Al2O3 influence; slag–iron interfacial co-reduction generates Fe3P intermediates; Fe3P incorporation into the iron matrix establishes stable solid solutions.

Graphical Abstract

1. Introduction

As the world’s largest steel producer, China faces critical supply chain vulnerabilities due to over 80% reliance on imported iron ore, threatening its steel industry’s sustainability [1,2]. Developing domestic low-grade iron resources thus becomes strategically imperative. High-phosphorus oolitic iron ore (HPOIO) constitutes a substantial reserve, with proven global deposits exceeding 20 billion tonnes—over 7 billion tonnes located in China [3,4,5]. Conventional mineral processing techniques, however, suffer from extended flowsheets, high operational costs, inefficient dephosphorization, and carbon-intensive operations, resulting in chronic underutilization of these resources.
Characterized as a refractory low-grade material, HPOIO typically contains 20%–50% Fe and 0.1%–1.5% P [6,7]. Its concentric oolitic texture incorporates hematite, quartz, chlorite, and apatite. Complex mineral intergrowth and ultrafine dissemination (0–3 μm) impede conventional beneficiation, while elevated phosphorus levels prevent production of high-grade, low-phosphorus iron concentrates [8,9]. Systematic development of alternative techniques—including conventional processing, magnetizing roasting-magnetic separation (MR-MS), direct reduction-magnetic separation (DR-MS), and hydrometallurgical methods—aims to overcome these limitations.
Current research on high-phosphorus oolitic iron ore (HPOIO) predominantly focuses on iron enrichment and phosphorus removal, aiming to produce low-phosphorus iron raw materials suitable for conventional blast furnace processes. However, micron-scale mineral dissemination in HPOIO presents dual challenges: existing grinding technologies fail to achieve effective mineral liberation, while excessively fine particles substantially increase grinding energy consumption. Furthermore, such ultrafine particles promote severe slime coating during separation, collectively resulting in inefficient beneficiation [10,11]. These factors constrain traditional mineral processing through high grinding energy demands, low iron recovery, and inadequate dephosphorization, ultimately preventing production of high-grade, low-phosphorus iron concentrates.
Although the MR-MS process enhances mineral processability by converting hematite to magnetite or γ-hematite via reduction roasting, it merely induces phase transformation without disrupting the original oolitic texture. Consequently, iron minerals remain intricately intergrown with phosphorus-bearing phases, yielding concentrates with persistently elevated phosphorus content [12,13]. In contrast, DR-MS employs high-temperature roasting to dismantle the oolitic structure and facilitate metallic iron aggregation. This enables secondary phase reconstruction between metallic and slag phases, permitting subsequent production of low-phosphorus iron powder through grinding and magnetic separation [14,15]. Nevertheless, DR-MS implementation requires roasting temperatures exceeding 1200 °C alongside substantial calcium/sodium-based dephosphorization reagents to inhibit phosphorus migration into metallic iron, imposing stringent limitations due to prohibitive energy consumption and production costs.
Hydrometallurgical methods (acid/bioleaching) selectively dissolve phosphorus minerals from HPOIO to prepare low-phosphorus iron products. However, these approaches encounter significant challenges, including excessive reagent consumption, protracted microbial cultivation periods, low processing efficiency, and acidic effluent pollution [16,17]. Collectively, these constraints compromise both technical feasibility and economic viability.
While phosphorus detrimentally compromises steel plasticity, impact toughness, weldability, and low-temperature performance, it delivers unique functional value in specialized materials [18,19]. High-phosphorus weathering steel Q295GNH (0.08%–0.15% P) forms dense protective layers through homogeneous dissolution, exhibiting 2–8 times greater atmospheric corrosion resistance than carbon steel while providing high strength, superior cold formability, and exceptional impact fatigue resistance for marine and automotive applications [20,21]. High-phosphorus cast iron (0.35%–0.65% P) demonstrates 1–3 times higher wear resistance than standard gray iron, serving ideally for machine tool guides and piston rings [22,23]. Medium- (0.7%–1.0% P) and high-phosphorus (2.0%–2.5% P) railway brake blocks leverage phosphorus-enhanced tribological properties for braking system applications [24,25], collectively validating phosphorus’s critical functional role in advanced materials.
This study innovatively proposes a pre-reduction and smelting separation (PR-SS) process that directly produces multifunctional Fe–P alloy from HPOIO. By circumventing conventional dephosphorization steps, PR-SS enables simultaneous iron and phosphorus enrichment, establishing an efficient pathway for strategic high-value utilization of HPOIO. Compared to traditional blast furnace ironmaking, this electric furnace-based process significantly reduces carbon emissions by eliminating the need for coke combustion. The compact process flow further enhances operational efficiency, utilizing coal, biocoke, or conventional coke as reductors without requiring sintering or coking operations. Consequently, PR-SS achieves substantially lower production and capital investment costs, representing a promising short-process alternative for Fe–P alloy synthesis.

2. Materials and Methods

2.1. Materials

The detailed chemical compositions of the two ores used in this experiment are presented in Table 1. The oolitic magnetite concentrate (OMC) was upgraded from the raw HPOIO ore via magnetizing roasting in a rotary kiln (850 °C, 15 min, 3% coal addition) followed by wet ball milling (95% of product <0.045 mm) and wet drum magnetic separation (1000 Gs magnetic field intensity). The raw HPOIO ore exhibited a total iron (TFe) grade of only 35.04%. After beneficiation, the TFe grade of OMC increased to 52.06%, yet it still contained high levels of SiO2 (16.32%), Al2O3 (9.86%), and P (0.37%). Due to its low iron grade and elevated phosphorus content, this ore is unsuitable for conventional Blast Furnace-Basic Oxygen Furnace (BF-BOF) processes and fails to meet industrial requirements. Consequently, exploring efficient non-blast furnace ironmaking technologies is critically imperative.
The laser particle size distribution characteristics of the OMC are illustrated in Figure 1a. The cumulative particle sizes D50 and D90 were measured as 17.891 μm and 48.511 μm, respectively, which means that the particle size of OMC was very fine. Figure 1b,c present the XRD patterns and SEM-EDS results of OMC, respectively, with Table 2 summarizing the point-scan EDS data. Microstructural analysis reveals that the ore primarily consists of magnetite ((Fe, Al)3O4), hematite ((Fe, Al)2O3), amorphous silicate (4FeO·Al2O3·3SiO2), quartz (SiO2), and fluorapatite (Ca5(PO4)3F). Magnetite is intergrown with silicates, quartz, and fluorapatite, forming intricate concentric banded structures. This configuration encapsulates iron minerals within layered matrices with ultrafine particle sizes (predominantly <5 μm), rendering effective liberation of magnetite monomers challenging. Critically, the iron-bearing silicate gangue phase (4FeO·Al2O3·3SiO2) is inevitably discarded into tailings during conventional beneficiation. As magnetic separation, gravity concentration, and flotation cannot recover iron from this phase, the overall iron recovery in the final concentrate is substantially reduced. Furthermore, extensive isomorphous substitution of aluminum in iron mineral lattices impedes Fe-Al separation. Figure 1d illustrates the oolitic texture and phosphorus distribution in HPOIO: hematite is tightly encapsulated by gangue minerals, forming characteristic ooids, with coarse-grained phosphorus minerals (0–10 μm) concentrated at ooid rims while fine-grained phosphorus phases predominantly reside in ooid cores. These mineralogical characteristics demonstrate significant challenges for mineral liberation and separation via traditional beneficiation approaches.
This study employed coal and coke as reductants for the pre-reduction and smelting separation stages, respectively, both pulverized to <5 mm before use. As detailed in Table 3, the fixed carbon (FCad) contents were 51.55% for coal and 82.81% for coke, with ash (Aad) contents of 9.44% and 14.81%, volatile matter (Vad) of 31.90% and 2.25%, and air-dried moisture (Mad) of 7.11% and 0.13%, respectively. These proximate analysis parameters meet the technical specifications for carbonaceous reductants in the PR-SS process, while the substantially low levels of detrimental impurities (S, P) comply with industrial production standards.

2.2. Experimental Methods

Figure 2 delineates the experimental procedure of the PR-SS process. Initially, OMC powder was compacted into cylindrical briquettes (Φ10 mm × 14 mm) using a PC-12 manual pellet press. A 50 g briquette was homogeneously blended with coal and loaded into a 200 mL alumina crucible for pre-reduction roasting at 1050 °C for 2 h with a C/Fe mass ratio of 2.0 in a KSL-1200X-J muffle furnace (Kejing Instrument, Hefei, China). Following pre-reduction, the briquette was mixed with coke at predetermined ratios in a dedicated alumina crucible and subjected to smelting separation in a VSF-type electric resistance furnace (atmosphere-controllable chamber furnace). Notably, the entire smelting separation process was conducted under an argon atmosphere. Upon reaching the designated smelting duration, the crucible was extracted, cooled to ambient temperature in argon, and manually disintegrated to segregate slag and metallic iron products. The metallization degree of pre-reduced briquettes was calculated using Equation (1), while iron and phosphorus recoveries in the smelting metal phase were determined via Equations (2) and (3), respectively.
φ = M F e T F e × 100 %
where φ is the metallization rate and MFe and TFe are the metallic iron and total iron contents in the sample, respectively.
θ = m × C M × R × 100 %
μ = m × D M × K × 100 %
where θ and μ represent the Fe recovery and P recovery (%), respectively; m represents the weight of the metallic iron (g); C and D represent the Fe grade and P grade of the metallic iron (%), respectively; M represents the weight of the feed (g); and R and K represent the Fe grade and P grade of the feed (%), respectively.

2.3. Analysis Methods

The chemical composition of the samples was analyzed via an X-ray fluorescence spectrometer (ZSX Primus IV, Rigaku, Saitama, Japan) and an inductively coupled plasma spectrometer (iCAP RPplus, Thermo Fisher Scientific, Waltham, MA, USA). A laser particle size analyzer (MS3000, Malvern, Worcestershire, UK) was used to determine the particle size and specific surface area of the sample. The mineral composition of each sample was determined through an advanced X-ray diffractometer (XRD) (D8 DISCOVER, BRUKER, Ettlingen, Germany). Optical microscopy (DMI5000 M, Leica, Karlsruhe, Germany) and field emission scanning electron microscopy (SEM) (MIRA3, Tescan, Brno, Czech Republic) were utilized to analyze the microstructure and crystal-chemical properties of the minerals. An electron probe X-ray microanalyzer (EPMA) (1720T, SHIMADZU, Kyoto, Japan) was used to measure the microelement composition of the sample. X-ray photoelectron spectroscopy (XPS) (K-Alpha, Thermo Fisher Scientific, Waltham, MA, USA) was employed to measure the elemental composition and valence states of the samples. ImageJ 1.44Plus (National Institutes of Health, Bethesda, MD, USA) was used to calculate the grain size of the crystals, and FactSage 8.3 (Thermfact/CRCT, Montreal, QC, Canada) was used to calculate the thermodynamic reactions and equilibrium phase diagrams.

3. Pre-Reduction and Smelting Separation Process

3.1. Preparation of Pre-Reduced Briquettes

Based on the optimized pre-reduction process parameters obtained in previous experiments, the optimal thermal conditions were determined as follows: C/Fe mass ratio of 2 and reduction temperature at 1050 °C for 2 h. Under these conditions, the pre-reduced briquettes were successfully prepared. Figure 3a shows that the original cylindrical morphology of the calcined briquettes was maintained without significant agglomeration, confirming the favorable operability and stability of the prereduction process. Figure 3b reveals that micron-sized metallic iron particles are uniformly dispersed within the briquettes, with no distinct boundaries between the slag and iron phases, indicating the inherent difficulty in achieving efficient phase-iron separation through only prereduction. Table 4 reported that pre-reduced briquettes possessed an iron grade of 64.19%, a phosphorus content of 0.39%, and a metallization rate of 74.21%, collectively validating the typical characteristics of HPOIO as a refractory low-grade iron source.

3.2. Effect of the Smelting Temperature

The effects of smelting temperature on slag–metal separation efficiency and Fe/P enrichment in the metallic phase at 5% coke ratio and 30 min duration are presented in Figure 4 and Figure 5. Experimental results demonstrate that at 1450 °C, slag and metal phases coexist with an indistinct separation interface. When the temperature exceeds 1500 °C, a well-defined interphase boundary forms, characterized by consolidated ellipsoidal metallic phases and inclusion-free slag. This phenomenon is attributed to reduced slag viscosity and diminished slag–metal interfacial tension at elevated temperatures, which synergistically enhance melt fluidity—the fundamental mechanism governing efficient phase separation [26,27].
As depicted in Figure 5a, elevating temperature from 1500 °C to 1600 °C maintained the total iron (TFe) grade in the metallic phase above 98%, while iron recovery exhibited a marginal decrease yet remained >96%. Figure 5b reveals that phosphorus content stabilized at ~0.7% with P recovery consistently exceeding 98%, indicating predominant phosphorus partitioning into the metallic phase. Through a comprehensive evaluation of energy consumption and slag–metal separation efficiency, 1550 °C was identified as the optimal smelting temperature, achieving 99.12% iron grade, 98.38% iron recovery, 0.72% phosphorus content, and 99.18% phosphorus recovery.

3.3. Effect of Coke Dosage

Figure 6 and Figure 7 illustrate the effects of coke dosage on slag–metal separation efficiency and the enrichment behavior of iron and phosphorus in the metallic phase at a smelting temperature of 1550 °C and a holding time of 30 min. As the coke dosage increased from 0% to 5%, the metallic phase formed a compact ellipsoidal morphology with distinct phase interfaces and no entrained iron particles within the slag phase. In contrast, at a coke dosage of 10%, the metallic phase existed as discrete particles that failed to coalesce. Further increasing the dosage to 15% resulted in mutual entrainment of iron particles and slag, accompanied by significant blurring of the slag–metal interface. This phenomenon is attributed to the fixed carbon within the coke, which possesses a smelting point exceeding 3000 °C. Excessive coke addition increases slag viscosity and reduces fluidity, thus impairing phase separation efficiency [28,29].
Figure 7a demonstrates that increasing the coke ratio from 0% to 5% maintains the iron grade of the metallic phase near 99%, while iron recovery substantially rises from 75.81% to 98.38%. Conversely, further elevation to 15% sharply reduces the iron grade to 89.77%, whereas iron recovery remains consistently >99%. This divergence stems from the dual functionality of coke during smelting: optimal coke loading facilitates iron mineral reduction to enhance recovery, while excess coke promotes carburization reactions with metallic iron, thereby diminishing purity [30,31,32]. Correspondingly, Figure 7b reveals a gradual decrease in phosphorus content within the metallic phase from 0.82% to 0.52% as the coke ratio increases from 0% to 15%. Phosphorus recovery exhibits a unimodal trend, peaking at 99.18% at the 5% coke ratio. This demonstrates that molten iron effectively absorbs phosphorus from the slag phase even without coke addition. Thus, a 5% coke ratio is established as optimal to ensure efficient slag–iron separation while simultaneously maximizing the recovery of both iron and phosphorus.

3.4. Effect of the Smelting Time

Figure 8 and Figure 9 depict the influence of smelting time on slag–metal separation efficiency and the enrichment behavior of iron and phosphorus in the metallic phase under conditions of 1550 °C and 5% coke dosage. At a smelting time of 10 min, macroscopic phase separation occurred between the slag and metal; however, distinct slag infiltration features were observed on the metal surface. Extending the smelting time to 20 min or longer resulted in a smooth and dense metal surface, with no discernible iron particles detected within the slag phase.
As shown in Figure 9a, increasing the smelting time from 10 to 40 min maintained the Fe grade in the metallic phase consistently above 98%. Meanwhile, the Fe recovery rate exhibited a marginal initial increase followed by a slight decrease, reaching its peak value of 99.73% at 20 min. Figure 9b indicates that the P content within the metallic phase remained essentially stable over the same smelting time range (10 to 40 min), while P recovery consistently exceeded 97%. Therefore, a smelting time of 20 min is identified as optimal, balancing slag–metal separation efficiency with energy consumption.
In summary, the optimal smelting conditions were identified as a temperature of 1550 °C, a duration of 20 min, and a coke dosage of 5%. Under these parameters, the metallic iron phase exhibited an Fe grade of 99.24%, an Fe recovery rate of 99.73%, a phosphorus content of 0.73%, and a phosphorus recovery rate of 99.15%.

3.5. Effect of the Metallization Rate of the Pre-Reduced Briquettes

Figure 10 and Figure 11 illustrate the influence of metallization degree on slag–metal separation efficiency and the enrichment behavior of iron and phosphorus in the metallic phase, demonstrating the adaptability of the PR-SS process to raw materials. With an increasing metallization degree from 57.31% to 74.21%, the metallic phase consistently exhibited a smooth ellipsoidal morphology. No visible iron particle inclusions were observed within the slag phase, and the slag–metal separation interface remained distinct. Concurrently, the iron recovery rate in the metallic iron phase increased progressively from 94.38% to 99.73%, while the Fe grade consistently exceeded 98%. Furthermore, the phosphorus content within the metallic phase remained stable, and phosphorus recovery exceeded 99% across the evaluated metallization range. These results indicate that an increased metallization degree enhances iron recovery in the metallic phase while exerting negligible influence on phosphorus recovery.

4. Characterization of the Smelting Sample

4.1. XRD

Figure 12 presents the XRD patterns of the metallic phase and slag phase obtained under optimal smelting conditions. The metallic phase exhibits distinct characteristic iron peaks, confirming its highly crystalline, single-phase metallic iron structure. In contrast, the slag phase displays a broad amorphous hump between 20° and 40° 2θ but also exhibits residual iron diffraction peaks. This indicates that the slag phase is predominantly an amorphous glassy phase, containing only trace amounts of crystalline iron phases.

4.2. EPMA

Figure 13 illustrates the EPMA elemental distribution maps of metallic and slag phases obtained under optimal smelting conditions, while Table 5 characterizes the chemical composition of the metallic phase. As shown in Figure 13a, a strong spatial correlation exists between iron and phosphorus within the metallic phase. Point analysis confirmed a phosphorus content of 0.715%, which aligns with the chemical analysis results, further confirming the enrichment of phosphorus within the iron matrix. Figure 13b demonstrates that phosphorus is largely absent from the vitreous regions of the slag phase. However, metallic iron particles entrapped within the slag exhibit significant phosphorus enrichment. This observation confirms the migration of phosphorus from the slag phase into the metallic iron phase.
The spatial distribution pattern of phosphorus validates its preferential enrichment within the metallic phase via an interfacial reaction and mass transfer mechanism. This process follows the transport pathway: slag phase → metal–slag interface → iron matrix, ultimately forming a homogeneous Fe–P solid solution.

5. Enrichment Mechanism of Phosphorus in Metallic Iron

5.1. Structural Destruction Mechanism of Ca3(PO4)2

Table 6 presents the formation energies of the investigated compounds. The formation energies of CaSiO3, CaAl2O4, and CaAl2(SiO4)2 are significantly lower than that of Ca3(PO4)2, indicating their superior thermodynamic stability at elevated temperatures [33,34]. Figure 14 elucidates the thermal decomposition mechanism of Ca3(PO4)2. Within the SiO2 and Al2O3 system, Ca2+ ions from Ca3(PO4)2 preferentially react with silico-aluminate oxides, forming the thermodynamically favored CaAl2(SiO4)2. This reaction-induced thermodynamic competition destabilizes the Ca3(PO4)2 lattice, liberating highly reactive P2O5. This highly active P2O5 component serves as the fundamental driving force for the subsequent migration and reduction of phosphorus.

5.2. Thermodynamics of Co-Reduction Between P2O5 and Iron Phase

Figure 15a demonstrates that under thermodynamic equilibrium conditions, the activity of P2O5 in the slag phase exhibits exponential growth with increasing temperature, indicating substantial generation of reactive P2O5 species at elevated temperatures. Table 7 and Figure 15b reveal that in systems containing solid carbon, P2O5 spontaneously undergoes co-reduction with iron phases (Fe2O3/Fe3O4/FeO/Fe) (ΔG < 0) to form Fe3P, thereby enabling efficient phosphorus reduction. Crucially, even in the absence of solid carbon, P2O5 participates in co-reduction with metallic iron to produce Fe3P and FeO, explaining phosphorus enrichment in the metallic phase during coke-free smelting processes.
Figure 16 illustrates the phosphorus migration pathway from slag to metallic phase, involving phosphate dissociation releasing reactive P2O5 at the slag–metal interface, followed by interfacial co-reduction of P2O5 with Fe to generate Fe3P, and subsequent dissolution equilibrium where Fe3P incorporates into the iron matrix to form a stable Fe–P solid solution. The activity of P2O5 in the slag phase governs phosphorus migration efficiency, as it constitutes the rate-determining step for this reaction sequence.

5.3. Fe–P Alloy Phase Analysis

To elucidate the phosphorus occurrence state in the metallic phase, Figure 17 illustrates the Fe–P binary alloy phase diagram [35,36,37]. Integrated with the sample composition from this study, a solidification phase transformation model was constructed to reconstruct the phase evolution pathway during the cooling of molten iron. As phosphorus-containing molten iron cools from 1550 °C to the liquidus temperature (1510 °C), the α-Fe phase preferentially precipitates from the liquid. Upon further cooling to 1380 °C, a eutectic reaction completes the transformation of the liquid phase into an α-Fe matrix. Subsequently, when cooled to the metastable region (800 °C), a eutectoid transformation induces the formation of Fe3P phases along α-Fe grain boundaries, ultimately yielding an α-Fe/Fe3P dual-phase structure. The solid solution characteristics of Fe3P in α-Fe enable the formation of Fe–P solid solutions via a vacancy-assisted diffusion mechanism.

6. Conclusions

This study conducted pre-reduction and smelting separation experiments using HPOIO, achieving high-efficiency phosphorus enrichment in the iron phase. Key conclusions are summarized as follows:
  • Process optimization reveals that under appropriate conditions (1550 °C, 20 min, 5% coke ratio), the metallic phase exhibits an iron grade of 99.24%, iron recovery of 99.73%, phosphorus content of 0.73%, and phosphorus recovery of 99.15%.
  • EPMA analysis demonstrates spatial correlation between iron and phosphorus in the metallic phase, while phosphorus signals remain undetected in the vitreous slag. This confirms preferential phosphorus enrichment in the metallic phase via interfacial mass transfer, following the migration pathway “slag phase → metal interface → iron matrix” to form homogeneous Fe–P solid solutions.
  • The phosphorus migration mechanism involves three stages: Initially, apatite lattice disintegration releases highly active P2O5 under the influence of SiO2 and Al2O3 components. Subsequently, co-reduction occurs at the slag–iron interface to form Fe3P intermediates. Ultimately, Fe3P incorporates into the iron matrix to establish stable Fe–P solid solutions. The activity of P2O5 in the slag governs migration efficiency, as it constitutes the rate-determining step in this reaction sequence.

Author Contributions

M.H.: conceptualization, investigation, and writing—original draft; D.Z.: validation, funding acquisition, and writing—review and editing; J.P.: methodology, resources, and supervision; Z.G.: validation and data curation; C.Y.: methodology; S.L.: project administration and funding acquisition; W.C.: Visualization. All authors have read and agreed to the published version of the manuscript.

Funding

This research was funded by the National Key Research and Development Program of China (Nos. 2023YFC3903900 and 2023YFC3903904), the National Natural Science Foundation Youth Foundation of China (No. 52404356), and the China Baowu Low Carbon Metallurgy Innovation Foundation (No. BWLCF202216), and The APC was funded by National Key Research and Development Program of China (Nos. 2023YFC3903900 and 2023YFC3903904).

Data Availability Statement

The raw data supporting the conclusions of this article will be made available by the authors on request.

Conflicts of Interest

The authors declare no conflicts of interest.

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Figure 1. Characteristics of ores. (a) Laser size distribution of OMC; (b) XRD pattern of OMC; (c) SEM-EDS image of OMC; (d) SEM-EDS image of HPOIO.
Figure 1. Characteristics of ores. (a) Laser size distribution of OMC; (b) XRD pattern of OMC; (c) SEM-EDS image of OMC; (d) SEM-EDS image of HPOIO.
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Figure 2. Experimental flow diagram of the PR-SS process.
Figure 2. Experimental flow diagram of the PR-SS process.
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Figure 3. Morphology of pre-reduced briquettes. (a) Macroscopic morphology; (b) microscopic morphology.
Figure 3. Morphology of pre-reduced briquettes. (a) Macroscopic morphology; (b) microscopic morphology.
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Figure 4. Effect of the smelting temperature on the separation of slag and iron.
Figure 4. Effect of the smelting temperature on the separation of slag and iron.
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Figure 5. Effect of the smelting temperature on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
Figure 5. Effect of the smelting temperature on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
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Figure 6. Effect of coke dosage on the separation of slag and iron.
Figure 6. Effect of coke dosage on the separation of slag and iron.
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Figure 7. Effects of the coke dosage on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
Figure 7. Effects of the coke dosage on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
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Figure 8. Effect of smelting time on the separation of slag and iron.
Figure 8. Effect of smelting time on the separation of slag and iron.
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Figure 9. Effect of the smelting time on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
Figure 9. Effect of the smelting time on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
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Figure 10. Effect of the metallization rate on the separation of slag and iron.
Figure 10. Effect of the metallization rate on the separation of slag and iron.
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Figure 11. Effects of the metallization rate on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
Figure 11. Effects of the metallization rate on the composition of the metallic phase. (a) Iron index; (b) phosphorus index.
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Figure 12. XRD patterns of the iron sample and slag sample.
Figure 12. XRD patterns of the iron sample and slag sample.
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Figure 13. EPMA of the iron sample and slag sample. (a) Iron sample; (b) slag sample.
Figure 13. EPMA of the iron sample and slag sample. (a) Iron sample; (b) slag sample.
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Figure 14. Structural destruction mechanism of Ca3(PO4)2.
Figure 14. Structural destruction mechanism of Ca3(PO4)2.
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Figure 15. Thermodynamic analysis of P2O5. (a) P2O5 activity in slag; (b) Gibbs free energy change (∆G).
Figure 15. Thermodynamic analysis of P2O5. (a) P2O5 activity in slag; (b) Gibbs free energy change (∆G).
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Figure 16. Diagram of phosphorus migration from the slag phase to the metal phase.
Figure 16. Diagram of phosphorus migration from the slag phase to the metal phase.
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Figure 17. Fe-P alloy phase diagram.
Figure 17. Fe-P alloy phase diagram.
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Table 1. Chemical composition of the ores (wt%).
Table 1. Chemical composition of the ores (wt%).
CompositionTFeFeOSiO2Al2O3CaOMgOK2ONa2OSPF
HPOIO35.0419.4425.4610.561.921.360.840.0430.370.430.10
OMC52.0626.5316.329.860.531.150.190.0210.110.370.07
Table 2. EDS spectrum analysis in Figure 1.
Table 2. EDS spectrum analysis in Figure 1.
SpectrumAt%Phase
OFMgAlSiPCaFe
139.7360.27SiO2
241.033.413.0452.52(Fe, Al)3O4
341.672.0015.4014.3826.554FeO∙Al2O3∙3SiO2
450.443.4815.9328.44Ca5(PO4)3F
Table 3. Industrial analysis of the reductants (wt%).
Table 3. Industrial analysis of the reductants (wt%).
TypeFCadMadAadVadSP
Coal51.557.119.4431.900.0020.001
Coke82.810.1314.812.250.0010.001
Table 4. Chemical composition of pre-reduced briquettes.
Table 4. Chemical composition of pre-reduced briquettes.
CompositionsTFeMFeSiO2Al2O3CaOMgOK2ONa2OSPF
Percentage (%)64.1947.6419.2811.490.581.210.210.0220.080.390.08
Table 5. Chemical composition of metallic phase sample.
Table 5. Chemical composition of metallic phase sample.
CompositionsTFeCSiO2Al2O3CaOMgOK2ONa2OSPF
Percentage (%)99.240.010.0010.0010.0010.0020.0010.0020.0020.730.001
Table 6. Thermodynamic stability of several phases.
Table 6. Thermodynamic stability of several phases.
TypesCa3(PO4)2CaSiO3CaAl2O4CaAl2(SiO4)2
Formation Energy (eV/atom)−3.326−3.448−3.454−3.426
Table 7. Possible reactions of [P2O5].
Table 7. Possible reactions of [P2O5].
Chemical ReactionEquation
P2O5 + 3Fe2O3 + 14C = 2Fe3P + 14CO(g)(4)
P2O5 + 2Fe3O4 + 13C = 2Fe3P + 13CO(g)(5)
P2O5 + 6FeO + 11C = 2Fe3P + 11CO(g)(6)
P2O5 + 6Fe + 5C = 2Fe3P + 5CO(g)(7)
P2O5 + 11Fe = 2Fe3P + 5FeO(g)(8)
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Hu, M.; Zhu, D.; Pan, J.; Guo, Z.; Yang, C.; Li, S.; Cao, W. Fe–P Alloy Production from High-Phosphorus Oolitic Iron Ore via Efficient Pre-Reduction and Smelting Separation. Minerals 2025, 15, 778. https://doi.org/10.3390/min15080778

AMA Style

Hu M, Zhu D, Pan J, Guo Z, Yang C, Li S, Cao W. Fe–P Alloy Production from High-Phosphorus Oolitic Iron Ore via Efficient Pre-Reduction and Smelting Separation. Minerals. 2025; 15(8):778. https://doi.org/10.3390/min15080778

Chicago/Turabian Style

Hu, Mengjie, Deqing Zhu, Jian Pan, Zhengqi Guo, Congcong Yang, Siwei Li, and Wen Cao. 2025. "Fe–P Alloy Production from High-Phosphorus Oolitic Iron Ore via Efficient Pre-Reduction and Smelting Separation" Minerals 15, no. 8: 778. https://doi.org/10.3390/min15080778

APA Style

Hu, M., Zhu, D., Pan, J., Guo, Z., Yang, C., Li, S., & Cao, W. (2025). Fe–P Alloy Production from High-Phosphorus Oolitic Iron Ore via Efficient Pre-Reduction and Smelting Separation. Minerals, 15(8), 778. https://doi.org/10.3390/min15080778

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