A method for recovering a target metal from iron or steel slag using at least one of a carbon thermal reduction process and a high-temperature wet smelting process.

Abstract

To provide pyro-hydrometallurgical methods to economically and environmentally recover a target metal from iron slag or steel slag.SOLUTION: For instance, the method enables subjecting an iron or steel slag feed to acid-baking with an acid to produce a dried mixture comprising at least one soluble metal salts, then subjecting the dried mixture to water leaching to make an aqueous solution comprising an aqueous leachate rich in the target metal and solid residues and subsequently separating the aqueous leachate rich in the target metal from the solid residues. This acid-baking water-leaching method facilitates efficient recovery of target metal compared to conventional methods.SELECTED DRAWING: Figure 1

Description

[Technical Field]

[0001] Related applications This application claims priority under applicable law to U.S. Provisional Application No. 62 / 824.588, filed on 27 March 2019, the contents of which are incorporated in their entirety by reference for all purposes.

[0002] The technical field generally relates to methods for recovering at least one valuable element from metallurgical slags such as iron slag and steel slag, more specifically, to methods for recovering at least one target metal from iron slag and steel slag using at least one of a carbon-thermal reduction process and a high-temperature wet smelting process. [Background technology]

[0003] As the importance of valorizing waste is increasingly emphasized in order to establish a circular economy, technosphere mining (i.e., the extraction of strategic materials from alternative secondary resources such as industrial waste residues) is gaining attention. For example, alternative secondary resources may include metallurgical slags such as iron-based slags, including iron slag and steel slag generated by the steel manufacturing or steel refining industries during the separation of molten iron or molten steel from impurities in ironmaking or steelmaking furnaces.

[0004] Approximately 15 to 20 million tons of steel slag are produced annually in the United States alone. Disposing of steel slag poses environmental risks. In fact, about 15% to 40% of steel slag is stored in steelmaking plants and then sent to landfills.

[0005] Iron and steel slags contain significant amounts of valuable materials, including titanium, niobium, platinum group metals (PGMs), gold, silver, and rare earth elements (REEs). For example, electric arc furnace (EAF) slag is an important potential source of many valuable metals, including but not limited to iron, manganese, magnesium, chromium, niobium, and aluminum.

[0006] These valuable materials can be recovered and given high value through extraction metallurgy. As a result, the extraction of valuable materials helps conserve natural resources and preserve landfill sites. It also means increased profits for the steel industry in an environmentally sustainable way.

[0007] Traditionally, the extraction of Nb, Ti, and REE from primary and secondary sources has been carried out by wet smelting methods such as direct acid leaching (Valighazvini et al., Industrial & Engineering Chemistry Research 52, no. 4 (2013): 1723-1730; El-Hussaini et al., Hydrometallurgy 64, no. 3 (2002): 219-229; and Kim et al., Minerals 6, no. 3 (2016): 63).

[0008] However, using conventional wet smelting methods can present several challenges. For example, conventional wet smelting methods require several pretreatment steps. Major concerns associated with conventional wet smelting methods include the consumption of large amounts of acids, bases, and organic solvents, as well as the production of large amounts of hazardous waste. In another example, Zheng et al. prepared titanium dioxide from Ti-supported electric furnace slag using the NH4HF2-HF leaching and hydrolysis process (Zheng et al., Journal of Hazardous Materials 344 (2018): 490-498). However, the use of hydrofluoric acid-based leaching agents has a significant negative impact on the environment.

[0009] Processes based on combined high-temperature wet smelting methods, such as acid calcination-leaching processes, are alternatives to direct leaching. Using high-temperature wet smelting, the metal-supported sample is mixed with concentrated acid, calcined in a furnace at an elevated temperature, and then leached at ambient temperature. Compared to traditional wet smelting methods, this process results in increased extraction efficiency with shorter residence times, smaller leaching volume, and significantly reduced acid consumption.

[0010] There are few reports on the extraction of valuable materials by high-temperature wet smelting methods. For example, Wu et al. proposed roasting Ta-Nb ore with concentrated sulfuric acid and leaching the calcined sample with dilute sulfuric acid (Wu et al., In Advanced Materials Research, vol. 997, pp. 651-654. Trans Tech Publications Ltd, 2014). Wu et al. achieved high leaching efficiency, but there are several drawbacks associated with acid leaching and high leaching temperatures. In another example, Qu et al. proposed extracting titanium from titanium-supported electric arc furnace slag by acid calcination and leaching with water at elevated temperatures of 50°C to 80°C (Qu et al., Journal of Materials Science Research Vol. 5, No. 4 (2016)). Qu et al. achieved only a moderate leaching efficiency of 84.3% and used high leaching temperatures. Achieved.

[0011] Carbon thermal reduction processes are also used to reduce metal oxides and thereby produce metals. For example, using this type of process, iron oxides are reduced by carbonaceous materials such as coal, coke, and natural gas to extract the metal.

[0012] Therefore, a process is needed to overcome one or more of the drawbacks encountered in conventional technosphere mining processes for iron slag and steel slag. [Overview of the Initiative]

[0013] According to a first aspect, the present technology relates to a method for recovering a target metal from iron or steel slag, wherein the method is: The process involves mixing iron or steel slag particles with acid in an acid-to-iron or steel slag particle mass ratio ranging from approximately 0.5 to approximately 5 to produce a mixture. The process involves heating the mixture at a temperature of approximately 100°C to approximately 600°C to remove excess water and acid, thereby producing a dry mixture containing pyrolysis gas and at least one soluble metal salt. The process involves adding water to obtain a density in the range of approximately 50 g / L to approximately 250 g / L to leach the dry mixture and produce an aqueous solution containing the aforementioned target metal-rich aqueous leachate and solid residue. The steps include separating the aqueous leachate rich in the aforementioned target metal from the solid residue, and Includes.

[0014] In one embodiment, the target metal is at least one of titanium, niobium, manganese, chromium, iron, scandium, neodymium, yttrium, lanthanum, cerium, samarium, gadolinium, dysprosium, praseodymium, europium, terbium, erbium, calcium, magnesium, aluminum, copper, silicon, ruthenium, rhodium, palladium, osmium, iridium, and platinum.

[0015] In another embodiment, the method further includes grinding the iron or steel slag particles before mixing. In one example, the grinding is carried out by ball milling.

[0016] In another embodiment, the method further includes classifying and separating iron or steel slag particles into fractions based on their size. In one example, the iron or steel slag particles have a size of less than about 200 mesh. In another example, the iron or steel slag particles have diameters ranging from about 1 μm to about 150 μm, or about 1 μm to about 140 μm, or about 1 μm to about 130 μm, or about 1 μm to about 120 μm, or about 1 μm to about 110 μm, or about 1 μm to about 100 μm, or about 1 μm to about 90 μm, or about 5 μm to about 90 μm.

[0017] In another embodiment, the method further includes drying the iron or steel slag particles before mixing. In one example, drying is carried out at a temperature in the range of about 50°C to about 80°C. In another example, drying is carried out for a period in the range of about 12 hours to about 24 hours.

[0018] In another embodiment, the acid is selected from the group consisting of sulfuric acid, hydrochloric acid, nitric acid, and mixtures of at least two of them. In one example, the acid includes sulfuric acid, and the soluble metal salt includes at least one soluble metal sulfate.

[0019] In another embodiment, the mass ratio of acid to iron or steel slag particles is in the range of about 1 to about 3, or about 2 to about 3, or about 1 to about 1.75.

[0020] In another embodiment, firing is carried out at a temperature of approximately 100°C to approximately 500°C, or approximately 200°C to approximately 600°C, or approximately 200°C to approximately 500°C, or approximately 200°C to approximately 400°C, or approximately 300°C to approximately 400°C.

[0021] In another embodiment, firing is carried out for a period of time ranging from approximately 30 minutes to approximately 240 minutes, or from approximately 30 minutes to approximately 120 minutes, or from approximately 30 minutes to approximately 60 minutes.

[0022] In another embodiment, the method further includes recycling the pyrolysis gas. In one example, recycling the pyrolysis gas includes reusing the pyrolysis gas in a mixing step.

[0023] In another embodiment, the leaching step is performed for a period of time ranging from about 30 minutes to about 360 minutes, or from about 120 minutes to about 360 minutes, or from about 180 minutes to about 360 minutes.

[0024] In another embodiment, the density of the dry mixture is in the range of about 100 g / L to about 200 g / L, or about 125 g / L to about 200 g / L, or about 150 g / L to about 200 g / L.

[0025] In another embodiment, the method further includes stirring an aqueous solution containing the aforementioned target metal-rich aqueous leachate and solid residue. In one example, the stirring is performed at a stirring speed of about 150 rpm to about 650 rpm.

[0026] In another embodiment, the separation involves filtering an aqueous solution containing the aforementioned target metal-rich aqueous leachate and solid residue.

[0027] In another embodiment, the method further includes purifying the aqueous leachate rich in the aforementioned target metal. In one example, the purification step is carried out by at least one of selective precipitation, solvent extraction, and ion exchange.

[0028] In another aspect, the present technology relates to a method for recovering niobium from iron or steel slag, wherein the method is: The process involves mixing iron or steel slag particles with acid in an acid-to-iron or steel slag particle mass ratio ranging from approximately 2.5 to approximately 3.5 to produce a mixture. The process involves heating the mixture at a temperature of approximately 375°C to 425°C to remove excess water and acid, thereby producing a dry mixture containing pyrolysis gas and at least one soluble niobium salt. The process involves leaching a dry mixture by adding water to obtain a density in the range of approximately 175 g / L to approximately 225 g / L, thereby producing an aqueous solution containing a niobium-rich aqueous leachate and solid residue. The steps include adding water to leach the dry mixture to obtain a density ranging from approximately 50 g / L to approximately 70 g / L, thereby producing a scandium-rich leachate aqueous solution and an aqueous solution containing solid residue, The process includes the step of separating a niobium-rich aqueous leachate from the solid residue.

[0029] In one embodiment, the acid is selected from the group consisting of sulfuric acid, hydrochloric acid, nitric acid, and mixtures of at least two of them. In one example, the acid includes sulfuric acid, and the soluble niobium salt includes at least one soluble niobium sulfate.

[0030] In another embodiment, the mass ratio of acid to iron or steel slag particles is about 3.

[0031] In another embodiment, the firing is carried out at a temperature of approximately 400°C.

[0032] In another embodiment, the density of the dry mixture is approximately 200 g / L.

[0033] In another embodiment, firing is carried out for a period ranging from approximately 30 minutes to approximately 240 minutes. In one example, firing is carried out for a period of approximately 120 minutes.

[0034] In another embodiment, the method further includes stirring an aqueous solution containing a niobium-rich aqueous leachate and solid residue at a stirring speed of about 150 to about 600 rpm. In one example, the stirring speed is about 150 rpm.

[0035] In another embodiment, the present technology relates to a method for recovering titanium from iron or steel slag, wherein the method is The process involves mixing iron or steel slag particles with acid in an acid-to-iron or steel slag particle mass ratio ranging from approximately 2 to approximately 3 to produce a mixture. The process involves heating the mixture at a temperature of approximately 200°C to approximately 400°C to remove excess water and acid, thereby producing a dry mixture containing pyrolysis gas and at least one soluble titanium salt. The process involves leaching a dry mixture by adding water to obtain a density in the range of approximately 50 g / L to approximately 200 g / L, thereby producing an aqueous solution containing a titanium-rich aqueous leachate and solid residue. A step of separating a titanium-rich aqueous leachate from the solid residue. Includes.

[0036] In one embodiment, the acid is selected from the group consisting of sulfuric acid, hydrochloric acid, nitric acid, and mixtures of at least two of them. In one example, the acid includes sulfuric acid, and the soluble titanium salt includes at least one soluble titanium sulfate.

[0037] In another embodiment, firing is performed for a period ranging from approximately 30 minutes to approximately 120 minutes, or from approximately 30 minutes to approximately 90 minutes.

[0038] In another embodiment, the density of the dry mixture is in the range of about 60 g / L to about 200 g / L.

[0039] In another embodiment, the method further includes stirring an aqueous solution containing a titanium-rich aqueous leachate and solid residue at a stirring speed in the range of about 150 rpm to about 550 rpm. In one example, the stirring speed is in the range of about 200 rpm to about 550 rpm.

[0040] In another embodiment, the steel slag is electric arc furnace slag, the acid-to-steel slag particle mass ratio is about 3, the calcination is carried out at a temperature of about 400°C, and the density of the dry mixture is about 200 g / L. In one example, the calcination is carried out for a period of about 120 minutes. In another example, the method further includes stirring an aqueous solution containing a titanium-rich aqueous leachate and solid residue at a stirring speed of about 150 rpm.

[0041] In another embodiment, the iron slag is blast furnace slag, the acid-to-iron slag particle mass ratio is about 2, the calcination is carried out at a temperature of about 200°C, and the density of the dry mixture is about 62 g / L. In one example, the calcination is carried out for a period of about 90 minutes. In another example, the method further includes stirring an aqueous solution containing a titanium-rich aqueous leachate and solid residue at a stirring speed of about 600 rpm.

[0042] In another aspect, the present technology relates to a method for recovering scandium from iron or steel slag, wherein the method is The process involves mixing iron or steel slag particles with acid in an acid-to-iron or steel slag particle mass ratio ranging from approximately 1.5 to approximately 2.5 to produce a mixture. The process involves heating the mixture at a temperature of approximately 175°C to 225°C to remove excess water and acid, thereby producing a dry mixture containing pyrolysis gas and at least one soluble scandium salt. The process involves leaching a dry mixture by adding water to obtain a density in the range of approximately 50 g / L to approximately 70 g / L, thereby producing an aqueous solution containing a scandium-rich aqueous leachate and solid residue. The steps include separating a scandium-rich aqueous leachate from the solid residue and Includes.

[0043] In one embodiment, the acid is selected from the group consisting of sulfuric acid, hydrochloric acid, nitric acid, and mixtures of at least two of them. In one example, the acid includes sulfuric acid, and the soluble scandium salt includes at least one soluble scandium sulfate.

[0044] In another embodiment, the mass ratio of acid to iron or steel slag particles is about 2.

[0045] In another embodiment, firing is carried out at a temperature of approximately 200°C.

[0046] In another embodiment, the density of the dry mixture is approximately 60 g / L.

[0047] In another embodiment, firing is carried out for a period ranging from approximately 60 minutes to approximately 120 minutes. In one example, firing is carried out for a period of approximately 90 minutes.

[0048] In another embodiment, the method further comprises stirring an aqueous solution containing a scandium-rich aqueous leachate and a solid residue at a stirring speed of about 600 rpm.

[0049] In another embodiment, the present technology relates to a method for recovering neodymium from iron or steel slag, wherein the method is The process involves mixing iron or steel slag particles with acid in an acid-to-iron or steel slag particle mass ratio ranging from approximately 1.5 to approximately 2.5 to produce a mixture. The process involves heating the mixture at a temperature of approximately 175°C to 225°C to remove excess water and acid, thereby producing a dry mixture containing pyrolysis gas and at least one soluble neodymium salt. The process involves leaching a dry mixture by adding water to obtain a density in the range of approximately 50 g / L to approximately 70 g / L, thereby producing an aqueous solution containing a neodymium-rich aqueous leachate and solid residue. A step of separating a neodymium-rich aqueous leachate from the solid residue. Includes.

[0050] In one embodiment, the acid is selected from the group consisting of sulfuric acid, hydrochloric acid, nitric acid, and mixtures of at least two of them. In one example, the acid includes sulfuric acid, and the soluble neodymium salt includes at least one soluble neodymium sulfate.

[0051] In another embodiment, the mass ratio of acid to iron or steel slag particles is about 2.

[0052] In another embodiment, firing is carried out at a temperature of approximately 200°C.

[0053] In another embodiment, the density of the dry mixture is approximately 60 g / L.

[0054] In another embodiment, firing is performed for a period ranging from approximately 30 minutes to approximately 60 minutes. In one example, firing is performed for a period of approximately 30 minutes.

[0055] In another embodiment, the method further includes stirring an aqueous solution containing a neodymium-rich aqueous leachate and a solid residue at a stirring speed of about 600 rpm.

[0056] In another embodiment, the present technology relates to a method for recovering a target metal from iron or steel slag, wherein the method is: A step of mixing iron or steel slag particles, at least one reducing agent with a reducing agent-to-iron or steel slag particle mass ratio in the range of about 0.06 to about 0.12, and at least one flux with a flux-to-iron or steel slag particle mass ratio in the range of 0 to about 0.1 to produce a mixture, The process involves smelting the mixture at a temperature of approximately 1300°C to 1800°C to form a metallic phase and a slag layer containing the target element, The steps include separating the metallic phase containing the target element from the slag phase to produce a metallic phase containing the target element and a slag layer, and Includes.

[0057] In one embodiment, the target metal is at least one of iron, manganese, chromium, and niobium. In one example, the target metal is iron.

[0058] In another embodiment, the iron or steel slag is electric arc furnace slag.

[0059] In another embodiment, the reducing agent includes a carbon source. In one example, the carbon source is selected from the group consisting of metallurgical coal, charcoal, petroleum coke, petroleum coke, natural gas, and at least two combinations thereof. In one important example, the carbon source is lignite.

[0060] In another embodiment, the mass ratio of reducing agent to iron or steel slag particles is about 0.06, about 0.09, or about 0.12.

[0061] In another embodiment, the flux is selected from the group consisting of silica, alumina, and mixtures thereof. In one important example, the flux is alumina. In another important example, the flux is silica. In yet another important example, the flux is a mixture of silica and alumina.

[0062] In another embodiment, the mass ratio of reducing agent to iron or steel slag particles is about 0.05 or about 0.1.

[0063] In another embodiment, the method further includes grinding the iron or steel slag particles before mixing. In one example, the grinding is carried out using at least one of a jaw crusher and a disc mill.

[0064] In another embodiment, the method further includes classifying and separating iron or steel slag particles into fractions based on their size.

[0065] In another embodiment, the iron or steel slag particles have a size of less than about 200 mesh.

[0066] In another embodiment, the method further includes drying the iron or steel slag particles before mixing. In one example, drying is carried out at a temperature of at least about 50°C. In another example, drying is carried out at a temperature in the range of about 50°C to about 80°C. In yet another example, drying is carried out for a period of at least about 24 hours.

[0067] In another embodiment, the method further includes pelletizing the mixture prior to the smelting step.

[0068] In another embodiment, the separation step is carried out by a mechanical separation method. In one example, the separation step is carried out manually.

[0069] In another embodiment, the method further includes grinding the slag phase to obtain slag particles.

[0070] In another embodiment, the method further includes classifying and separating the slag particles into fractions based on their size.

[0071] In another embodiment, the method further includes reducing the content of the target metal in the slag phase to produce a target metal-depleted slag. In one example, reducing the content of the target metal in the slag phase is done by magnetic separation. For example, magnetic separation is performed using Davis® tubes.

[0072] In another embodiment, the smelting step is carried out at temperatures in the range of approximately 1300°C to approximately 1700°C, or approximately 1300°C to approximately 1600°C, or approximately 1400°C to approximately 1800°C, or approximately 1400°C to approximately 1700°C, or approximately 1400°C to approximately 1600°C, or approximately 1500°C to approximately 1800°C, or approximately 1500°C to approximately 1700°C, or approximately 1500°C to approximately 1600°C. In one example, the smelting step is carried out at temperatures in the range of approximately 1500°C to approximately 1600°C.

[0073] In another embodiment, iron or steel slag is a byproduct of an ironmaking or steelmaking process. In one example, the ironmaking or steelmaking process provides thermal energy to the iron or steel slag. For example, the temperature of the iron or steel slag particles is raised before the mixing step.

[0074] In another embodiment, the iron or steel slag particles and the reducing agent undergo a redox reaction that releases chemical energy. For example, a redox reaction generates heat as a byproduct (exothermic reaction). Alternatively, the iron or steel slag particles and the reducing agent undergo a redox reaction that does not release chemical energy.

[0075] In another embodiment, the method further includes subjecting the target metal depletion slag to a high-temperature wet smelting process to recover a second target metal. In one example, the high-temperature wet smelting process is a method as described herein. [Brief explanation of the drawing]

[0076] [Figure 1] This is a flowchart of a process for recovering valuable elements from iron slag and steel slag according to one embodiment. [Figure 2] This is a schematic diagram of a process for recovering valuable elements from iron slag and steel slag according to one embodiment. [Figure 3] This is a flowchart of a process for recovering valuable elements from iron slag and steel slag according to another embodiment. [Figure 4] This is a graph of the particle size distribution of pulverized electric arc furnace (EAF) slag, as described in Example 1(a). [Figure 5] The results of characterization of the pulverized EAF slag, as described in Example 1(a), are shown in (a) and (b) graphs of the inductively coupled plasma atomic emission spectroscopy (ICP-OES) results showing the elemental composition, (c) the X-ray diffraction pattern, and (d) scanning electron microscope (SEM) images of the slag particles (scale bar represents 5 μm) and energy-dispersive X-ray spectroscopy (EDS) elemental mapping of the slag particles. [Figure 6] The results of the extraction efficiency of several elements at different factor levels, as described in Example 1(d), are shown, with (a) a graph of the calcination temperature, (b) the acid-to-slag mass ratio, (c) the calcination time, (d) the slag-to-water density, and (e) the stirring rate. [Figure 7] Graphs (a) and (b) show the rate test results for various elements at firing temperatures of 200°C and 400°C, respectively, as described in Example 1(d), and graph (c) shows the extraction efficiency of various elements at different acid-to-slag mass ratios, including additional test results at ratios of 1.25, 1.50, and 1.75. [Figure 8] The graph shows the factor effect coefficients of the empirical sampling model, as described in Example 1(e), with (a) the result for Nd, (b) the result for Ti, (c) the result for Fe, (d) the result for Ca, (e) the result for Mn, (f) the result for Mg, (g) the result for Al, and (h) the result for Cr. The inset graph shows the correlation between the results predicted by the empirical model and the experimental results. [Figure 9] The results of characterization of EAF slag before and after the acid calcination process, as described in Example 1(g), are shown in (a) the XRD (powder X-ray diffraction) diffractograms of EAF slag before and after the acid calcination process, and (b) to (e) show SEM images of various slag particle samples prepared under various acid calcination conditions. [Figure 10] As described in Example 1(h), (a) shows the XRD diffractogram of calcium titanate before and after the acid calcination process, (b) is a diagram of the 3D structure of calcium titanate, and (c) to (m) are SEM images of calcium titanate before and after the acid calcination process and EDS elemental mapping for various elements. [Figure 11] This is a graph of the particle size distribution of pulverized blast furnace (BF) slag, as described in Example 2(a). [Figure 12]The results of characterization of the pulverized BF slag, as described in Example 2(a), are shown in (a) and (b) graphs of the ICP-OES and inductively coupled plasma mass spectrometry (ICP-MS) results showing the elemental composition, (c) is a powder X-ray diffractogram, and (d) is a SEM image and EDS elemental mapping of the BF slag particles. [Figure 13] The graphs show the rate test results for rare earth elements and base metals under other intermediate operating parameter conditions, such as those described in Example 2(d), at (a) a level 1 firing temperature condition (200°C) and (b) a level 1 firing temperature condition (400°C). [Figure 14] The graph shows the factor effect coefficients of the empirical sampling model, as described in Example 2(e), with (a) the result for Sc, (b) the result for Nd, (c) the result for Ca, (d) the result for Al, (e) the result for Mg, and (f) the result for Fe. The inset graph shows the correlation between the results predicted by the empirical model and the experimental results. [Figure 15] This is a graph showing the predicted and actual extraction efficiencies for Sc, Nd, Al, Mg, and Fe, as described in Example 2(f). [Figure 16] The graph shows the factor effect coefficients of the empirical sampling model, as described in Example 4(e), with (a) the result for Fe, (b) the result for Nb, (c) the result for Cr, (d) the result for Mn, and (e) the result for Ti. The inset graph shows the correlation between the results predicted by the empirical model and the experimental results. [Figure 17] The results of electron beam microanalyzer (EPMA) phase mapping of the metallic phase obtained after carbonothermal reduction, as described in Example 4(e), are shown. [Figure 18] The characterization results of Fe-depleted slag, as described in Example 4(g), are shown in (a) a graph of the ICP-OES results showing the elemental composition, (b) a powder X-ray diffractogram, (c) the results of EPMA phase mapping, and (d) a secondary electron image (scale bar represents 10 μm) and EDS elemental mapping of Fe-depleted slag particles. [Figure 19]The graph shows the factor effect coefficients of the empirical sampling model, as described in Example 4(h), with (a) showing the results for Ti, (b) for Fe, (c) for Ca, (d) for Mn, (e) for Mg, (f) for Al, (g) for Cr, and (h) for Sr. The inset graph shows the correlation between the results predicted by the empirical model and the experimental results. [Figure 20] The following two-dimensional contour plots show the interaction effect of two main factors of the acid calcination water leaching process on the average extraction efficiency, as described in Example 4(h): (a) shows the results for calcination temperature (X1) and acid-to-slag mass ratio (X2), (b) shows the results for calcination temperature (X1) and calcination time (X3), (c) shows the results for calcination temperature (X1) and water-to-slag ratio (X4), (d) shows the results for acid-to-slag mass ratio (X2) and calcination time (X3), (e) shows the results for acid-to-slag mass ratio (X2) and water-to-slag ratio (X4), and (f) shows the results for calcination time (X3) and water-to-slag ratio (X4). [Figure 21] This is a graph showing the predicted and actual extraction efficiencies for Ti, Fe, Ca, Mn, Mg, Al, Cr, and Sr, as described in Example 4(i). [Modes for carrying out the invention]

[0077] Detailed explanation The following detailed description and examples are illustrative and should not be construed as further limiting the scope of the invention. On the contrary, it is intended to cover all alternatives, modifications, and equivalents that may be included as defined herein. The purpose, advantages, and other features of the method will become clearer and better understood by reading the following non-limiting description and referring to the accompanying drawings.

[0078] All technical and scientific terms and expressions used herein have the same definitions as those commonly understood by those skilled in the art when relating to this art. However, for clarification, definitions of some terms and expressions used herein are provided below.

[0079] When the term "about" is used herein, it means approximately, in the region of, or roughly. When the term "about" is used in relation to a number, it modifies the number by, for example, a variation of plus or minus 10% of its nominal value. The term may also take into account the probability of random error in experimental measurements, such as rounding of numbers or instrument limitations. When a range of values ​​is referred to in this application, the lower and upper limits of the range are always included in the definition unless otherwise indicated. When a range of values ​​is referred to in this application, it is therefore intended to include all intermediate and partial ranges as well as the individual values ​​included within the range.

[0080] The expression "particle size" in this specification refers to the distribution of the particle size. x It is described by , where the value dx is such that x% of the particle is d x Represents diameters less than d. 10 The value indicates that 10% of all particles have a particle diameter smaller than that value. 90 The value is such that 90% of all particles have a particle diameter smaller than that value. Therefore, d 50 The value represents the median particle diameter, meaning that 50% of all particles are larger than this diameter, and 50% are smaller than this diameter.

[0081] Where the term “equilibrium” is used herein, it refers to a steady state in which an ongoing process attempts to change a variable, but the variable in question has no observable (or net) effect on the properties of the system.

[0082] When the term "pore" is used herein, it refers to the space found within a particle, i.e., the void space (intraparticle pore) within a porous particle.

[0083] As used herein, the term “dry” refers to a process by which at least some of the water is removed from the material being dried. As used herein, the term “dried” or its synonym “dried” material defines, unless otherwise specified, the total moisture content of the aforementioned material being 5.0 wt.% or less based on the total weight of the dried material.

[0084] As used herein, the term “soaking” or its synonym “soaking” refers to the decomposition or dissolution of a sample in a strong acid such as concentrated sulfuric acid (H2SO4), hydrochloric acid (HCl), and nitric acid (HNO3) in order to obtain a paste-like sample and then raise the temperature to obtain a powder-like sample.

[0085] Throughout the following description, it is worth noting that when the article “a” is used to introduce an element, it does not mean “just one,” but rather “one or more.” Where the specification states that a step, component, feature, or characteristic “may,” “might,” “can,” or “could,” it should be understood that that particular component, feature, or characteristic does not have to be included in all substitutes. Where the term “comprising” or its synonyms “including” or “having” is used herein, it does not exclude other elements. For the purposes of the present invention, the expression “consisting of” is considered a preferred embodiment of the term “comprising.” Where a group is defined below to include at least a certain number of embodiments, it should also be understood that a group consisting of only these embodiments is preferably disclosed.

[0086] The various methods described herein relate to the recovery of at least one valuable element from iron-based slag (e.g., iron slag and steel slag) by high-temperature wet smelting processes. For example, the high-temperature wet smelting process includes both acid calcination and water leaching, hereafter referred to as acid calcination water leaching (ABWL).

[0087] More specifically, this technology relates to a method for recovering at least one valuable element from iron slag or steel slag. The valuable element is the target metal. For example, the target metal may be selected from transition metals, platinum group metals, metalloids, post-transition metals, alkaline earth metals, and / or rare earth metals. Non-limiting examples of target metals include titanium, niobium, manganese, chromium, iron, scandium, neodymium, yttrium, lanthanum, cerium, samarium, gadolinium, dysprosium, praseodymium, europium, terbium, erbium, calcium, magnesium, aluminum, copper, silicon, ruthenium, rhodium, palladium, osmium, iridium, and platinum. Preferably, the target metal is at least one of titanium, niobium, manganese, chromium, iron, scandium, neodymium, yttrium, lanthanum, cerium, samarium, gadolinium, dysprosium, praseodymium, europium, terbium, erbium, calcium, magnesium, aluminum, copper, and silicon. According to important variations, the target metal is at least one of titanium, niobium, scandium, and neodymium.

[0088] In some embodiments, two or more valuable elements may be co-recovered (or co-extracted). Alternatively, one valuable element may be selectively recovered (or extracted) while preventing the co-recovery (or co-extraction) of other elements.

[0089] Iron or steel slag may contain at least one component selected from the group consisting of silicon dioxide (SiO2), calcium oxide (CaO), iron(III) oxide (Fe2O3), iron(II) oxide (FeO), aluminum oxide (Al2O3), magnesium oxide (MgO), manganese(II) oxide (MnO), phosphorus pentoxide (P2O5), sulfur, and at least two combinations thereof. For example, iron or steel slag mainly contains SiO2, CaO, Fe2O3, FeO, Al2O3, MgO, MnO, P2O5, and / or sulfur.

[0090] In some cases, steel slag is steelmaking furnace slag. For example, steelmaking furnace slag can be basic oxygen converter (BOF) slag, electric arc furnace (EAF) slag, or ladle slag. In one significant variation, steelmaking furnace slag is either EAF slag or BOF slag. For example, the components of EAF slag may include, but are not limited to, FeO, Fe2O3, MnO, CaO, and at least two combinations thereof. For example, EAF slag is mainly composed of FeO, Fe2O3, MnO, and / or CaO.

[0091] In other cases, iron slag is blast furnace (BF) slag (ironmaking slag). The components of BF slag may include Al2O3, MgO, and sulfur. For example, BF slag is mainly composed of Al2O3, MgO, and / or sulfur.

[0092] In some embodiments, the iron or steel slag further comprises at least one other component, including but not limited to magnesium, calcium, vanadium, tungsten, copper, lead, zinc, REE, uranium, and thorium.

[0093] For a more detailed understanding of the disclosure, first refer to Figure 1, which provides a flowchart of a method for recovering at least one valuable element from iron or steel slag, according to possible embodiments.

[0094] As shown in Figure 1, a method for recovering at least one valuable element (or target metal) from slag includes the step of mixing iron or steel slag particles with an acid to produce a mixture.

[0095] Iron or steel slag particles and acid are mixed together in an acid-to-iron or steel slag particle mass ratio ranging from about 0.5 to about 5. For example, the acid-to-iron or steel slag particle mass ratio may range from about 1 to about 4, or about 1 to about 3.5, or about 1 to about 3, or about 2 to about 3, or about 1 to about 2, or about 1 to about 1.75. According to significant variations, the acid-to-iron or steel slag particle mass ratio is in the range of about 2 to about 3.

[0096] For example, the acid may be selected from the group consisting of sulfuric acid, hydrochloric acid, nitric acid, and mixtures of at least two of them, where applicable. In one significant variation, the acid includes sulfuric acid.

[0097] Acids can be concentrated acids. For example, acids can have concentrations ranging from approximately 95% to approximately 99.999%.

[0098] Referring to Figure 1, the method also includes the step of calcining the mixture and immersing it in hot water to remove excess water and acid, thereby producing a dry mixture containing pyrolysis gas and at least one soluble metal salt.

[0099] Firing is carried out at temperatures in the range of approximately 100°C to approximately 600°C. For example, firing may be carried out at temperatures in the range of approximately 200°C to approximately 600°C, or approximately 200°C to approximately 550°C, or approximately 200°C to approximately 500°C, or approximately 200°C to approximately 450°C, or approximately 200°C to approximately 400°C, or approximately 200°C to approximately 300°C, or approximately 300°C to approximately 400°C. According to significant variations, firing is carried out at temperatures in the range of approximately 200°C to approximately 400°C.

[0100] In some embodiments, firing is performed for a period of at least 30 minutes. For example, firing may be performed for a period ranging from about 30 minutes to about 240 minutes, or about 30 minutes to about 180 minutes, or about 30 minutes to about 120 minutes, or about 30 minutes to about 90 minutes, or about 30 minutes to about 60 minutes. In one significant variation, firing is performed for a period of about 120 minutes.

[0101] In cases where the acid contains sulfuric acid, the soluble metal salt contains at least one soluble metal sulfate.

[0102] In cases where the acid contains sulfuric acid, the pyrolysis gas may contain at least one of sulfur trioxide and sulfur dioxide. As shown in Figure 1, the pyrolysis gas produced during the calcination step can be recycled. For example, the pyrolysis gas can be reused in the mixing step, thereby reducing chemical costs and increasing the environmental sustainability of the method.

[0103] The mixing and firing steps may be performed sequentially, simultaneously, or with partial overlap of each other's timing. In some embodiments, the mixing and firing steps are performed sequentially, with the mixing step taking place before the firing step.

[0104] In some embodiments, the mixing and firing steps are carried out in a high-temperature processing apparatus, such as a rotary kiln.

[0105] In some embodiments, the iron or steel slag is a byproduct of an ironmaking or steelmaking process (not shown in Figure 1). For example, the ironmaking or steelmaking process provides thermal energy to the iron or steel slag, and the iron or steel slag particles may be at a temperature that has risen before the mixing and firing steps, thereby reducing the energy input required to obtain the temperature needed for the firing steps.

[0106] Referring to Figure 1, the method also includes leaching the dry mixture by adding water to produce an aqueous solution containing the aforementioned target metal-rich aqueous leachate and solid residue.

[0107] Water is added to obtain a water-to-dry mixture density in the range of approximately 50 g / L to approximately 250 g / L. In at least one embodiment, the water-to-dry mixture density is in the range of approximately 60 g / L to approximately 200 g / L, or approximately 75 g / L to approximately 200 g / L, or approximately 100 g / L to approximately 200 g / L, or approximately 125 g / L to approximately 200 g / L, or approximately 150 g / L to approximately 200 g / L, or approximately 55 g / L to approximately 200 g / L.

[0108] In some embodiments, the leaching step is performed at ambient temperature. In some embodiments, the leaching step is performed until the system reaches equilibrium. For example, the leaching step is performed for a period ranging from about 30 minutes to about 360 minutes, or from about 120 minutes to about 360 minutes, or from about 180 minutes to about 360 minutes.

[0109] Referring to Figure 1, the method also includes separating the aqueous leachate rich in the target metal from the solid residue.

[0110] In some embodiments, the separation includes filtering an aqueous solution containing the aforementioned target metal-rich aqueous leachate and solid residue. For example, the filtration may be at least one of the following: vacuum filtration, pressure filtration, etc. Alternatively, in significant variations, the separation includes a sedimentation step in which the solid residue settles at the bottom of the container containing the aforementioned target metal-rich aqueous leachate, forming a precipitate or slurry. For example, the sedimentation may be gravity sedimentation or centrifugal sedimentation.

[0111] As shown in Figure 1, the method may further include grinding the iron or steel slag to obtain iron or steel slag particles before mixing. According to a significant variation, the grinding is carried out by ball milling. For example, ball milling may activate the particles, thereby increasing the extraction efficiency. In some embodiments, the method further includes sorting and separating the iron or steel slag particles into fractions by size. For example, the slag particles may be sorted and classified into narrow-width fractions to obtain iron or steel slag particles having a size of less than about 200 mesh. In some embodiments, the iron or steel slag particles are substantially uniform in size. For example, iron or steel slag particles may have diameters ranging from about 1 μm to about 150 μm, or about 1 μm to about 140 μm, or about 1 μm to about 130 μm, or about 1 μm to about 120 μm, or about 1 μm to about 110 μm, or about 1 μm to about 100 μm, or about 1 μm to about 90 μm, or about 5 μm to about 90 μm. According to significant variations, iron or steel slag particles may have diameters ranging from about 1 μm to about 130 μm.

[0112] Referring to Figure 1, according to some embodiments, the method may further include drying the iron or steel slag particles before mixing. For example, drying may be carried out at a temperature of at least about 50°C. For example, drying may be carried out at a temperature in the range of about 50°C to about 80°C. For example, drying may be carried out for a period of at least about 12 hours. For example, drying may be carried out for a period in the range of about 12 hours to about 24 hours.

[0113] In some embodiments, the method further includes stirring an aqueous solution containing the aforementioned target metal-rich aqueous leachate and solid residue (not shown in Figure 1). Stirring and leaching may be performed sequentially, simultaneously, or with partial overlap of each other's time. In some embodiments, stirring and leaching are performed simultaneously. For example, stirring may be performed at a stirring speed in the range of about 150 rpm to about 650 rpm. For example, the stirring speed may be in the range of about 200 rpm to about 600 rpm.

[0114] As shown in Figure 1, excess acid may be recovered after the separation step. For example, the recovered acid may be reused in the mixing step, thereby reducing costs and increasing the environmental sustainability of the method.

[0115] In some embodiments, the method may further include purifying the aqueous leachate rich in the aforementioned target metal (not shown in Figure 1). For example, the purification step may be carried out by at least one of selective precipitation, solvent extraction, and ion exchange.

[0116] In at least one embodiment, the target metal is niobium. The mass ratio of iron or steel slag particles to acid is in the range of about 2.5 to about 3.5, preferably about 3. The calcination is carried out at a temperature of about 375°C to about 425°C, preferably about 400°C. The density of the dry mixture is in the range of about 175 g / L to about 225 g / L, preferably about 200 g / L. For example, the calcination may be carried out for a period of time in the range of about 30 minutes to about 240 minutes, or about 30 minutes to about 150 minutes, or about 30 minutes to about 120 minutes, preferably about 120 minutes. For example, the method may include stirring an aqueous solution containing a niobium-rich aqueous leachate and solid residue at a stirring speed in the range of about 150 rpm to 600 rpm, preferably about 150 rpm.

[0117] In some embodiments, the extraction efficiency of niobium may be at least about 90%. For example, the extraction efficiency of niobium may be in the range of at least about 92%, or at least about 95%, or about 92% to about 99%, or about 92% to about 98%, or about 92% to about 97%.

[0118] In at least one embodiment, the target metal is titanium. The mass ratio of acid to iron or steel slag particles is in the range of about 2 to about 3. The calcination is carried out at a temperature of about 200°C to about 400°C. The density of the dry mixture is in the range of about 50 g / L to about 200 g / L, or about 60 g / L to about 200 g / L. For example, the calcination may be carried out in the range of about 30 minutes to about 120 minutes, or about 30 minutes to about 90 minutes. For example, the method further includes stirring an aqueous solution containing a titanium-rich aqueous leachate and solid residue at a stirring speed in the range of about 150 rpm to about 550 rpm, or about 200 rpm to about 550 rpm. In one embodiment, the steel slag is electric arc furnace slag, the mass ratio of acid to steel slag particles is about 3, the calcination is carried out at a temperature of about 400°C, and the density of the dry mixture is about 200 g / L. The calcination is carried out for approximately 120 minutes, with a stirring speed of approximately 150 rpm. In one embodiment, the iron slag is BF slag, the mass ratio of acid to iron slag particles is approximately 2, the calcination is carried out at a temperature of approximately 200°C, and the density of the dry mixture is approximately 62 g / L. The calcination is carried out for approximately 90 minutes, with a stirring speed of approximately 600 rpm.

[0119] In some embodiments, the titanium extraction efficiency may be at least about 90%. For example, the titanium extraction efficiency may be at least about 92%, or at least about 95%, or at least about 97%.

[0120] In at least one embodiment, the target metal is scandium. The mass ratio of iron or steel slag particles to acid is in the range of about 1.5 to about 2.5, preferably about 2. The calcination is carried out at a temperature of about 175°C to about 225°C, preferably about 200°C. The density of the dry mixture is in the range of about 50 g / L to about 70 g / L, preferably about 60 g / L. The calcination is carried out for a period of about 60 minutes to about 120 minutes, preferably about 90 minutes. The method may further include stirring the aqueous solution containing the scandium-rich aqueous leachate and solid residue at a stirring speed in the range of about 550 rpm to about 650 rpm, preferably about 600 rpm.

[0121] In some embodiments, the scandium extraction efficiency may be at least about 70%. For example, the scandium extraction efficiency may be in the range of about 70% to about 98%, or about 70% to about 89%, or about 70% to about 85%, or about 75% to about 85%, or about 75% to about 82%.

[0122] In at least one embodiment, the target metal is neodymium. The mass ratio of iron or steel slag particles to acid is in the range of about 1.5 to about 2.5, preferably about 2. The calcination is carried out at a temperature of about 175°C to about 225°C, preferably about 200°C. The density of the dry mixture is in the range of about 50 g / L to about 70 g / L, preferably about 60 g / L. The calcination is carried out for a period of about 30 minutes to about 60 minutes, preferably about 30 minutes. The method may further include stirring the aqueous solution containing the neodymium-rich aqueous leachate and solid residue at a stirring speed in the range of about 550 rpm to about 650 rpm, preferably about 600 rpm. In some embodiments, the extraction efficiency of neodymium is less than about 60%, or less than about 40%.

[0123] In some embodiments, the extraction efficiency of iron can be at least about 85%. For example, the extraction efficiency of Fe can be at least about 90%, or at least about 91%. In some embodiments, the extraction efficiency of manganese can be at least about 90%. For example, the extraction efficiency of Mn can be at least about 95%, or at least about 97%. In some embodiments, the extraction efficiency of magnesium can be at least about 95%. For example, the extraction efficiency of Mg can be at least about 97%, or at least about 98%. In some embodiments, the extraction efficiency of aluminum can be at least about 58%. For example, the extraction efficiency of Al can be at least about 65%, or at least about 75%, or at least about 80%. In some embodiments, the extraction efficiency of chromium can be at least about 90%. For example, the extraction efficiency of Cr can be at least about 95%, or at least about 97%, or at least about 98%.

[0124] For a more detailed understanding of the disclosure, see Figure 2, which provides a schematic diagram of the ABWL recovery system 10 according to possible embodiments. The ABWL recovery system 10 includes a high-temperature treatment device 12 and a water leaching unit 14. According to one embodiment, the recovery system 10 may further include a sedimentation container 16.

[0125] In some embodiments, the high-temperature processing apparatus 12 is configured to receive a first feed stream 18 containing iron or steel slag directly from, for example, an industrial waste line (not shown in Figure 2), and a second feed stream 20 containing an acid, such as sulfuric acid. As shown in Figure 2, the high-temperature processing apparatus 12 is further configured to discharge a dry mixture 22. Referring to Figure 2, the water leaching unit 14 is configured to receive a third feed stream 24 containing the dry mixture 22 and water. According to one embodiment, the water leaching unit 14 may further include a stirring device 26. The water leaching unit 14 is further configured to discharge an aqueous solution 28 containing a target metal-rich aqueous leachate and solid residue. As shown in Figure 2, the settling container 16 is configured to receive the aqueous solution 28 containing the target metal-rich aqueous leachate and solid residue and to discharge solid residue 30 and a target metal-rich leachate 32.

[0126] This application also describes various methods for recovering at least one valuable element from iron or steel slag by a carbon-thermal reduction process. For example, the carbon-thermal reduction process can be used alone or in combination with a high-temperature wet smelting process.

[0127] In some embodiments, the carbon thermal reduction process may be used in combination with a high-temperature wet smelting process, such as the ABWL process described herein.

[0128] In some embodiments, the carbon-thermal reduction process may be carried out before the high-temperature wet smelting process.

[0129] More specifically, the technology also relates to a method for recovering at least one valuable element from iron or steel slag. The valuable element is the target metal. For example, the target metal may be selected from transition metals, metalloids, post-transition metals, alkaline earth metals, and / or rare earth metals. Non-limiting examples of target metals include iron, calcium, silicon, manganese, aluminum, chromium, strontium, copper, nickel, titanium, and niobium. According to significant variations, the target metal may be at least one of iron, manganese, chromium, niobium, magnesium, and aluminum. In some embodiments, the steel slag is EAF slag. For example, EAF slag may contain at least one of iron, manganese, chromium, magnesium, niobium, and aluminum.

[0130] In some embodiments, two or more valuable elements may be co-recovered (or co-extracted). Alternatively, one valuable element may be selectively recovered (or extracted) while preventing the co-recovery (or co-extraction) of other elements.

[0131] In some embodiments, at least one valuable element can be selectively extracted as a metallic phase from iron or steel slag by a carbonothermal reduction process. For example, at least one of iron, manganese, chromium, and niobium can be selectively extracted as a metallic phase from iron or steel slag by a carbonothermal reduction process. In one significant variation, iron is selectively extracted as a metallic phase.

[0132] In some embodiments, at least one valuable element may be selectively extracted as a metallic phase from iron or steel slag by a carbon-thermal reduction process, and at least one other valuable element may be extracted by a subsequent high-temperature wet smelting process. For example, at least one of magnesium and aluminum may be extracted by a subsequent high-temperature wet smelting process. In some examples, at least one of iron, manganese, chromium, and niobium may be selectively extracted as a metallic phase from iron or steel slag by a carbon-thermal reduction process, and at least one of magnesium and aluminum may be extracted by a subsequent high-temperature wet smelting process.

[0133] For example, a carbon-thermal reduction process may be used to substantially reduce the presence of at least one valuable element from iron or steel slag and to substantially increase the extraction of at least one target metal by a subsequent high-temperature wet smelting process.

[0134] For a more detailed understanding of the disclosure, refer next to Figure 3, which provides a flowchart of a method for recovering at least one valuable element from iron or steel slag, according to possible embodiments.

[0135] As shown in Figure 3, a method for recovering at least one valuable element (or target metal) from iron or steel slag includes the step of mixing iron or steel slag particles, at least one reducing agent, and optionally at least one flux.

[0136] Non-limiting examples of suitable reducing agents include carbon sources such as metallurgical coal, charcoal, petroleum coke, petroleum coke, natural gas, or a combination of at least two of these. In one significant variation, the reducing agent includes lignite.

[0137] In some embodiments, iron or steel slag particles and a reducing agent are mixed in a carbon-to-slag mass ratio of at least about 0.06 g of carbon per gram of iron or steel slag. For example, iron or steel slag particles and a reducing agent are mixed in a carbon-to-slag mass ratio ranging from about 0.06 to about 0.12 g of carbon per gram of iron or steel slag, including a limiting value. For example, iron or steel slag particles and a reducing agent are mixed in a carbon-to-slag mass ratio of about 0.06 g of carbon per gram of iron or steel slag, or about 0.09 g of carbon per gram of iron or steel slag, or about 0.12 g of carbon per gram of iron or steel slag.

[0138] Examples of suitable fluxes include, but are not limited to, silica (SiO2), alumina (Al2O3), and combinations of at least two of them. In one significant variation, the flux is silica. In another significant variation, the flux is alumina. In yet another significant variation, the flux includes both alumina and silica. For example, the flux may be selected for its ability to promote liquid phase formation or to remove chemical impurities from the metallic phase of iron or steel slag.

[0139] In some embodiments, iron or steel slag particles and flux are mixed in a flux-to-slag mass ratio ranging from 0 to about 0.1 g of flux per gram of iron or steel slag. For example, iron or steel slag particles and flux are mixed in a flux-to-slag mass ratio of 0 g of flux per gram of iron or steel slag, or about 0.05 g of flux per gram of iron or steel slag, or about 0.1 g of flux per gram of iron or steel slag.

[0140] As shown in Figure 3, the method may further include crushing the iron or steel slag before mixing to obtain iron or steel slag particles. For example, any suitable crushing method is intended. According to significant variations, the crushing is carried out using at least one of a jaw crusher and a disc mill. In some embodiments, the method further includes sorting and separating the iron or steel slag particles into fractions by size. For example, the slag particles may be sorted and classified into narrow-width fractions to obtain iron or steel slag particles having a size of less than about 200 mesh (about 74 μm). In some examples, the iron or steel slag particles are substantially uniform in size.

[0141] Referring to Figure 3, the method may further include drying the iron or steel slag particles before mixing in order to remove moisture content. For example, drying may be carried out at a temperature of at least about 50°C. For example, drying may be carried out at a temperature in the range of about 50°C to about 80°C. For example, drying may be carried out for a period of at least about 24 hours.

[0142] Referring to Figure 3, the method may further include a step of pelletizing a mixture containing iron or steel slag particles, a reducing agent, and a flux to produce pellets. For example, the pelletizing step may be performed to substantially increase the contact area between the reducing agent and the iron or steel slag particles, thereby increasing the reduction rate in the subsequent smelting step.

[0143] In some embodiments, the pelletizing step may be carried out by pressing the mixture into pellets using a press. Any suitable pelletizing method is intended. According to significant variations, pelletizing is carried out by pressing the mixture into pellets using an air / hydraulic press. In some examples, pelletizing is carried out at a pressure of about 250 MPa for about 3 minutes, and under these conditions, for example, the pellets have a diameter of about 28.6 mm.

[0144] Referring to Figure 3, the method also includes the step of smelting a mixture or pellet containing iron or steel slag particles, a reducing agent, and a flux to obtain the target metal as a slag phase and a metallic phase. The smelting step is carried out by applying heat to the mixture or pellet to extract the target metal. Any suitable means for applying heat is intended. For example, the smelting step is carried out using a melting or smelting furnace.

[0145] In some embodiments, the smelting step is carried out at a temperature in the range of about 1300°C to about 1800°C, with a heating rate of about 180°C per hour, a cooling rate of about 180°C per hour, and a holding time of about 1.5 hours. For example, the smelting step is carried out at a temperature in the range of about 1300°C to about 1700°C, or about 1300°C to about 1600°C, or about 1400°C to about 1800°C, or about 1400°C to about 1700°C, or about 1400°C to about 1600°C, or about 1500°C to about 1800°C, or about 1500°C to about 1700°C, or about 1500°C to about 1600°C, including a limit value. In one significant variation, the smelting step is carried out at a temperature in the range of about 1500°C to about 1600°C, including a limit value. In some embodiments, the smelting step may be carried out in an inert atmosphere, for example, an argon atmosphere. In one significant variation, the smelting step is carried out at a temperature of approximately 1500°C. In another significant variation, the smelting step is carried out at a temperature of approximately 1550°C. In yet another significant variation, the smelting step is carried out at a temperature of approximately 1600°C. For example, increasing the smelting temperature reduces the viscosity of the iron or steel slag, thereby improving the separation of the target metal as a metallic phase from the slag phase and resulting in increased extraction efficiency.

[0146] In some embodiments, the iron or steel slag is a byproduct of an ironmaking or steelmaking process (not shown in Figure 3). For example, the ironmaking or steelmaking process provides thermal energy to the iron or steel slag, and the iron or steel slag particles may be at a temperature that has risen before the mixing and smelting steps, thereby reducing the energy input required to obtain the temperature necessary for the smelting steps.

[0147] In some embodiments, the iron or steel slag particles and the reducing agent undergo a redox reaction that releases chemical energy. For example, the redox reaction generates heat as a byproduct (exothermic reaction). Alternatively, the iron or steel slag particles and the reducing agent undergo a redox reaction that does not release chemical energy. For example, if the redox reaction releases chemical energy, it may reduce the energy input required to obtain the temperature necessary for the smelting step.

[0148] As shown in Figure 3, the method also includes the step of separating the metallic phase (e.g., iron) from the slag phase. Any suitable means for separating the metallic phase from the slag phase is intended. For example, the metallic phase may be separated from the slag phase manually, for example, using any suitable mechanical separation method.

[0149] Referring to Figure 3, the method further includes the step of grinding the slag phase to obtain slag particles. For example, any suitable grinding method is intended. In some embodiments, the method further includes classifying and separating the slag particles into fractions by size. For example, the slag particles may be sorted and classified into narrow-width fractions to obtain slag particles having substantially reduced size. In some examples, the slag particles are substantially uniform in size.

[0150] Referring to Figure 3, the method may optionally include a step of further reducing the content of the metallic phase (e.g., iron) in the slag phase to produce a slag depleted of the target metal, such as iron-depleted slag.

[0151] In some embodiments, the step of further reducing the metallic phase content in the slag phase may be carried out by magnetic separation. Any magnetic separation method is intended. For example, magnetic separation may be carried out using Davis® tubes.

[0152] Referring to Figure 3, the method may further include subjecting a slag depleted of the target metal (e.g., iron-depleted slag) to a high-temperature wet smelting process to recover at least one valuable element from the slag. For example, the high-temperature wet smelting process may be an ABWL process as described herein.

[0153] The recovery of valuable elements from iron or steel slag can address the supply of strategic materials, the protection of natural resources, and related sustainability challenges, effectively revealing the hidden value of industrial waste. The technology described herein can significantly reduce the amount of waste to be disposed of, resulting in cost savings and environmental benefits. In addition, recycling the pyrolysis gases and waste leachates generated in the wet smelting process can further reduce chemical costs. This acid calcination water leaching method can facilitate the efficient recovery of the target metal compared to conventional methods. Another advantage of the technology described herein is that it may be possible to use the residual heat of the iron or steelmaking process in the acid calcination step, which also means reducing costs and benefiting the environment by advocating a more environmentally sustainable method. Thus, this method is applicable to adding value to iron or steel slag.

[0154] Examples The following non-limiting embodiments are illustrative and should not be construed as limiting the scope of the invention. These embodiments will be better understood with reference to the accompanying drawings.

[0155] Example 1: Recovery of valuable elements from EAF slag using ABWL (a) Preparation and characterization of EAF slag particles EAF slag samples were crushed and ground using a ball mill to obtain EAF slag particles. The EAF slag particles were then separated and classified into narrow-width fractions to obtain substantially uniform particle sizes of approximately 200 mesh. The separated EAF slag particles were then dried in a convection oven for more than 24 hours.

[0156] The EAF slag particles were characterized by X-ray diffraction (XRD), energy-dispersive spectroscopy attached to a scanning electron microscope (SEM-EDS), inductively coupled plasma mass spectrometry (ICP-MS), and inductively coupled plasma optical emission spectrometry (ICP-OES).

[0157] The particle size distribution of the crushed EAF slag was measured using a particle size analyzer. As shown in Figure 4, the median particle size was about 16.0 μm, the average particle size was about 25.3 μm, and D 90 was about 58.4 μm.

[0158] As shown in Figure 5, the elemental composition of the EAF slag was characterized by ICP-OES after aqua regia leaching (Figure 5(a) and (b)). The EAF slag sample contained about 0.05 wt.% Nb, about 0.19 wt.% Ti, about 24 wt.% Fe, about 17 wt.% Ca, about 6 wt.% Mn, about 6 wt.% Mg, about 3 wt.% Al, and about 2 wt.% Cr.

[0159] As shown in Figure 5(c), several oxide phases were detected, including wüstite Fe 0.944 O, hematite Fe2O3, brownmillerite Ca2((Fe 1.63 Al 0.37 )O5), spinel magnesioferrite Fe2MgO4, and larnite Ca2Si. The results indicated that the sample was mainly in oxide form, but other phases that were not identified due to the complexity of the sample peaks may have been present. Furthermore, the low concentrations of Nb and Ti hindered the detection of phases containing these elements.

[0160] The surface morphology and elemental mapping were examined using SEM-EDS as shown in Figure 5(d). The EDS results showed that Nb and Ti were concentrated in specific regions and were highly correlated with each other and with Ca. For example, Ti may exist as calcium titanate, and Nb may exist in the same compound by isomorphous substitution.

[0161] (b) Recovery of metals from EAF slag particles After drying, several samples were prepared by mixing 2 g of selected EAF slag particles prepared in Example 1(a) with concentrated (96 wt.%) H2SO4, and then calcining the mixture in furnaces at different temperatures from 200°C to 400°C. The acid-calcined samples were then leached in water at 25°C for 6 hours under magnetic stirring at different slag-to-water densities and stirring rates. The leached solutions were diluted using a Hamilton® Microlab 600® dual dilution and dispensing system and characterized by ICP-OES. The samples were characterized by XRD and SEM-EDS before and after acid calcination.

[0162] (c) Experimental design and empirical model building A systematic study was conducted to investigate the quantitative effects of five operating factors—namely, acid calcination temperature (X1), acid-to-slag particle mass ratio (X2), acid calcination time (X3), leached pulp density (or slag particle-to-water density) (X4), stirring speed (X5), and combinations of these factors—on the extraction efficiency of Nb, Ti, Fe, Ca, Mn, Mg, Al, and Cr.

[0163] For statistical analysis, the test factors were encoded into low (-1), medium (0), and high (+1) levels to allow for direct comparison of the relative effects of each factor based on the magnitude of the factor model coefficients.

[0164] Detailed information about the factor levels is shown in Table 1. The upper and lower limits of the factor levels were selected based on preliminary experiments, within the operating range where the system's response could be substantially linear. [Table 1]

[0165] The experimental matrix with these five operational factors was designed using a partially factorial design, one of the statistical methods used to evaluate the effect of each factor on the response and the interactions between the factors.

number

[0166] Additional tests were conducted with X2 values ​​of 1.25, 1.50, and 1.75 to examine the effect of X2, and velocity studies were performed for 7 hours at X1 = 200°C and 400°C. These experiments disrupted the balance of the design matrix and are therefore not included in the empirical model building dataset.

[0167] For building empirical models

number

[0168] Target element i(y i The experimental response was determined based on ICP-OES measurements of the concentration and the extraction efficiency calculated as outlined in Equation 1.

number

[0169] In the formula, C i This is the concentration of the analyte (target element) in the leachate, V i m0 is the volume of the leachate, and m0 is the weight of each target element present in the raw material sample used.

[0170] The effects of each studied factor and combinations of these factors on extraction efficiency were quantitatively evaluated by fitting experimental data, and an empirical model for each target element given by Equation 2 (below) was constructed to determine the extraction efficiency of the target element.

number

[0171] The model is fitted to the experimental data by linear multiple regression analysis (mLLSR) using the least squares method shown in Equation 3 (below), where,

number

number

[0172] Based on the two-tailed t-distribution shown in Equation 4 (below), the significance of each model parameter is expressed using a 95% confidence interval (CI). 95% This was verified by determining the following: Parameter distribution

number

number

[0173] Analysis of variance (ANOVA) was performed on the simplified model to evaluate the accuracy of the model's agreement. The coefficient of determination (R) was calculated. 2 ), measured analyte extraction efficiency (y i ) and predicted values

number

number

[0174] (d) Determination of operational parameters and evaluation of their effect on the extraction efficiency of Nb, Ti, Fe, Ca, and other elements. The effects of ABWL operation parameters on the extraction efficiency of Nb, Ti, Fe, Ca, and other elements were investigated with respect to X1, X2, X3, X4, and X5.

[0175] Table 3 below provides an overview of the operating conditions and the corresponding extraction efficiencies for Nb, Ti, Fe, Ca, Mn, Mg, Al, and Cr. In these trials, the maximum extraction efficiencies for Nb and Ti were 100%, while other elements such as Fe, Mn, and Mg were co-extracted with efficiencies exceeding 95%. [Table 3]

[0176] In Table 3, runs 17-21 are validation runs analyzed using standard stock solutions of Nb and Ti at a concentration of 100 mg / L, and standard stock solutions of Fe, Ca, Mn, Mg, Al, and Cr at a concentration of 1000 mg / L, supplied by Inorganic Ventures, Inc.

[0177] The effects of each factor (X1-X5) were determined, and the extraction efficiencies of each element at different factor levels are shown in Figure 6.

[0178] As shown in Figure 6(a), X1 has a significant positive effect on the extraction efficiency of Nb and Cr at a leaching time of 6 hours. At the same leaching time, the extraction efficiency of Ca is negatively affected by the increase in X1.

[0179] As seen in Figure 6(e), X5 has virtually no effect on the extraction efficiency of almost all elements except Ca, and for Ca, X5 has a small but positive effect. These results indicate that X5 at 150 rpm is sufficient for the system to reach chemical equilibrium and extract almost all elements from the EAF slag, but the secondary effect of this parameter should be considered as an accurate predictor of the maximum extraction efficiency.

[0180] As shown in Figures 7(a) and (b), water leaching rate experiments were performed on samples calcined at 200°C and 400°C for up to 7 hours. The results showed that the kinetics of the leaching process were slower for the sample prepared at X1=400°C compared to the sample prepared at X1=200°C, and that, given sufficient time to reach equilibrium, increasing X1 had a positive effect on the leaching efficiency of all elements except Ca, presumably because Ca precipitates as CaSO4 in the sulfuric acid medium, which has very low water solubility.

[0181] Among the operating factors considered, X2 showed the greatest positive effect on the extraction of most elements, which may be due to the fact that a higher acid-to-slag ratio results in more acid molecules reacting with the slag particle sample, effectively leading to a significant increase in extraction efficiency.

[0182] It should be noted that for Fe and Ti, X²=2 is sufficient to reach their respective maximum extraction efficiencies. Therefore, the results at the center point at factor level 0 (ratio 2) were similar to those at factor level +1 (ratio 3). Thus, additional experiments in the range of factor level -1 (ratio 1) to level 0 (ratio 2) are necessary to gain further insight into the effect of these factors on the extraction efficiencies of Fe and Ti. Additional tests were performed at acid-to-slag mass ratios of 1.25, 1.50, and 1.75, and the results are shown in Figure 7(c).

[0183] As shown in Figure 7(c), the extraction efficiencies of Nb and Ca have a linear relationship with X2, more specifically in the range of 1 to 3, while a secondary interaction can also be observed between X2 and Fe, Ti, and other elements (quadratic curvature).

[0184] (e) Building empirical models An empirical extraction model was constructed using mMLSR, as shown in Equation 2, to evaluate the relative effects of the principal experimental factors and aliased secondary interactions. Figure 8 shows the ordered factor effect coefficients of the empirical extraction model for each element.

[0185] As shown in Figure 8(a), the factors that have the greatest positive effect on Nb extraction are X1 and X3. The reason behind this observation is Nb 5+ However, Ti in the CaTiO3 phase 4+ This means that it may be replaced by Nb. 5+ The charge density of Ti 4+ Because the charge density is higher than that of , this substitution results in greater stability. Therefore, more energy is required for the sulfation reaction of Nb. It is also known that Nb oxides only begin to dissolve in the sulfuric acid medium at high temperatures. Another reason for the positive effects of X1 and X3 is that increasing X1 to 400°C results in the conversion of the (H3O)Fe(SO4)2 phase to the Fe2(SO4)3 phase, which releases H2SO4 and thereby increases the extraction efficiency of Nb. Increasing X3 increases the residence time and therefore has a positive effect on the extraction efficiency.

[0186] In addition to X1 and X3, Nb extraction is slightly positively affected by an increase in X2. Based on the Eh-pH diagram (or Pourbaix diagram, also known as the potential / pH diagram), Nb2O5 is soluble under the acidic conditions of the water leaching step, which explains the less significant effect of X2 on the Nb extraction efficiency.

[0187] In the case of Ti, the extraction efficiency increases significantly with increasing X2, suggesting that in the hot immersion reaction, the acid is the limiting reagent, and therefore, the more acid molecules available, the more Ti molecules are converted to soluble Ti sulfate (Figure 8(b)). As mentioned above, Ti can exist within the Ca-supported phase as CaTiO3 with a cubic crystal structure. Since Ti is at the center of the unit cell and surrounded by Ca atoms, Ca must first react with H2SO4 in order to extract Ti during the acid calcination process. This explains the positive effect of X2 on the Ti extraction efficiency. According to the Eh-pH diagram, Ti2(SO4)3 is soluble only at low pH, which further explains the positive effect of X2.

[0188] Other elements (Fe, Mn, Mg) exhibit extraction behavior similar to Ti, but the magnitude of each factor's effect differs slightly (Figures 8(c), (e), and (f)). X4 shows a very small or slight positive effect on the extraction efficiency of all elements except Ca (Figure 8(d)). X5 has a very small effect on the extraction efficiency of all elements except Ca, which can be expected because a positive effect of the stirring speed can only be observed when membrane diffusion is the rate-limiting step. Therefore, a stirring speed of 150 rpm is sufficient to extract almost all elements from the EAF slag.

[0189] X4 has a positive effect on the extraction efficiency of almost all elements. For this reason, 200 g / L can be considered sufficient to extract almost all elements from EAF slag.

[0190] Ca exhibits unique extraction behavior, where X4 has a negative effect and X2 has a positive effect. At high X4, more solid is available for leaching if sufficient acid is present for extraction, but the solubility of CaSO4 is low in water. Therefore, more Ca precipitates from the solution at high X4, and thus a decrease in extraction efficiency may be observed. The positive effect of X2 may be due to the fact that the solubility of CaSO4 increases with increasing H2SO4 concentration in the pH range considered in this example.

[0191] Detailed information on the primary and secondary effects on the extraction efficiency of Nb, Ti, Fe, Ca, Mn, Mg, Al, and Cr is shown in Table 4. [Table 4]

[0192] (f) Acid calcination water leaching process From this empirical extraction model, the optimized recovery conditions were determined to be high X1 (400°C, +1 level), high X2 (3, +1 level), high X3 (120 minutes, +1 level), high X4 (200 g / L, +1 level), and low X5 (150 rpm, +1 level). Experiments were conducted under the optimal conditions to verify the applicability of the empirical model. Table 5 shows the extraction efficiency predicted from the empirical model, along with the actual leaching efficiency. [Table 5]

[0193] As shown in Table 5, extractions of 98.7% Nb, 93.4% Ti, and over 97% Fe, Mn, and Mg were achieved under these conditions. Therefore, it was concluded that the empirical model constructed in this study can successfully predict optimal operating conditions and corresponding extraction efficiencies with high accuracy (absolute mean relative deviation (AARD) = 12.8%).

[0194] (g) Characterization of EAF slag before and after the acid calcination process The crystalline structure and surface morphology of EAF slag before and after acid calcination under different operating conditions were investigated by XRD and SEM-EDS (Figure 9). Figure 9(a) shows the XRD diffractograms of EAF slag particles before and after the acid calcination and water leaching processes. SEM-EDS images are shown in Figures 9(b) to (e), where (b) shows the raw EAF slag particles, (c) shows the sample prepared at 200°C, (d) shows the sample calcined at 400°C, and (e) shows the residue obtained after the water leaching step.

[0195] These samples were selected to study the effects of acid calcination, calcination temperature, and water leaching. The results of XRD and SEM are shown in Figure 9. In the EAF slag, Nb and Ti are present in trace amounts and are associated with Ca, as previously described with respect to Figure 5(d). Therefore, investigating the leaching mechanism of Nb and Ti requires observation of the mineralogical and morphological changes of the major element (i.e., Ca or Fe) phase during leaching in the acid calcination process.

[0196] To investigate the mineralogical changes of the Ca and Fe phases during acid calcination, samples calcined at 200°C and 400°C were characterized by XRD. As shown in Figure 9(a), the XRD results indicate that during the acid calcination process, the Fe main phase, namely wustite (FeO) and hematite (Fe2O3), as well as the Ca-supported phase (Ca2SiO4), are immersed and replaced by sulfate phases. At lower calcination temperatures (200°C), oxonium iron sulfate double salt ((H3O)Fe(SO4)2), zomolnokite (FeSO4·H2O), and anhydrite (CaSO4) are formed, while the iron phase is converted to anhydrous Fe2(SO4)3 at higher calcination temperatures. Based on the XRD results, the following reactions are expected to occur at 200°C: 4FeO (s) +8H2SO 4(aq) +O 2(g) →4(H3O)Fe(SO4) 2(s) +2H2O (g) [Formula 10] Ca2SiO 4(s) +2H2SO 4(aq) →2CaSO 4(s) +H4SiO4(aq) [Formula 11]

[0197] At higher firing temperatures (400°C), the FeSO4·H2O and (H3O)Fe(SO4)2 phases were converted to Fe2(SO4)3 according to the following reaction, as shown in Figure 9(a): 2(H3O)Fe(SO4) 2(s) →Fe2(SO4) 3(s) +H2SO 4(g) +2H2O (g) [Formula 12] 4FeSO4·H2O (s) +4H2SO 4(aq) →2Fe2(SO4) 3(s) +2SO 2(g) +8H2O (g) [Formula 13] 4FeSO4·H2O (s) +2H2SO 4(aq) +O 2(g) →2Fe2(SO4) 3(s) +6H2O (g) [Formula 14]

[0198] As seen in Equation 12, the conversion of (H3O)Fe(SO4)2 to Fe2(SO4)3 releases further H2SO4. This additional acid may participate in the sulfation reaction, thereby increasing the extraction efficiency of the target element. Unlike FeSO4·H2O and (H3O)Fe(SO4)2, which have a weak layered hydrogen bonding structure, Fe2(SO4)3 produced at 400°C has a cubic crystal structure, and therefore the leaching kinetics of the sample calcined at 400°C are slower than those of the sample calcined at 200°C, as shown in Figures 7(a) and (b).

[0199] As can be seen in Figure 9(a), the residue after the water leaching step consists mainly of CaSO4, which has low water solubility, resulting in low Ca extraction efficiency.

[0200] The microfluidic transport environment of the silicate-rich phase can sometimes lead to the polymerization of silica gel, which has an adverse effect on the leaching process. As shown in Figure 9(a), the main silicate phase (CaSiO4) in the EAF slag was completely leached during the acid calcination process, and the Si-bearing phase was not detected in the samples acid-calcined at both firing temperatures. As a result, the formation of silica gel was not detected during the water leaching step.

[0201] The morphological changes of the EAF slag before firing, after firing at 200 °C and 400 °C, and after water leaching were characterized using SEM. The EAF slag raw material shown in Figure 9(b) has a rough surface with an average particle size of 25.3 μm (Figure 4). As shown in Figure 9(c), when the slag sample was acid-calcined at a low temperature (200 °C), the morphology changed to large plate-like crystals, which was consistent with the six-sided flat plate morphology reported for the crystals of (H3O)Fe(SO4)2. In contrast, as shown in Figure 9(d), the slag sample fired at a high temperature (400 °C) had a cubic crystal structure, which was consistent with the structure of rhombohedral Fe2(SO4)3. Figure 9(e) shows the SEM image of the sample after water leaching, which is 100% CaSO4. These microscopic structure observations are consistent with the XRD spectra.

[0202] (h) Mechanistic study of EAF slag extraction ABWL experiments were conducted to extract Ti from CaTiO3 as a reference sample.

[0203] Figures 10(a) to (m) show the XRD and SEM-EDS results of the CaTiO3 samples before and after the acid calcination process. As shown in the XRD diffractogram of Figure 10(a), the CaTiO3 phase is mainly converted into four different phases, namely, anhydrite (CaSO4), titanium sulfate (Ti2(SO4)3), oxotitanium sulfate (TiOSO4), and titanium oxide (TiO2) after the acid calcination process. The change in the crystal structure may suggest that the following reactions occurred during the ABWL process. CaTiO3 + H2SO4 → CaSO4 + TiO2 + H2O [Equation 15] TiO2 + H2SO4 → TiOSO4 + H2O [Equation 16] 4TiOSO4 + 2H2SO4 → 2Ti2(SO4)3 + 2H2O + O2 [Equation 17]

[0204] When CaTiO3 reacts with H2SO4, the precipitation of anhydrite and the formation of TiO2 first occur, followed by the reaction between TiO2 and H2SO4, and it has been demonstrated that either Ti2(SO4)3 or TiOSO4 is formed.

[0205] As shown in Figure 10(b), CaTiO3 has a cubic crystal structure of space group Pm-3m, and Ti is at the center of the unit cell, surrounded by eight Ca atoms. Therefore, in order to extract Ti using the acid calcination process, Ca must first be sulfated by H2SO4 to form CaSO4 (Figure 8(b)).

[0206] Figures 10(c) to (m) show the SEM micrographs and EDS element mappings of the samples before and after acid calcination, and the results are in good agreement with the results of Ca and Ti obtained by XRD. In the sample before acid calcination, Ca and Ti were highly related to each other as CaTiO3, having a particle size of less than about 3 μm and a substantially smooth surface. However, after the acid calcination process, the particle size increased to about 20 - 30 μm, and the surface of the particles became significantly rough and porous. Furthermore, the results of element mapping showed that Ti and Ca were separated and concentrated in different regions after the acid calcination process. This implies that Ca was converted to CaSO4 which precipitated easily, while Ti2(SO4)3 and TiOSO4 particles were generated after the reaction between Ti and H2SO4 and aggregated to form larger particles. Although it is well known that TiOSO4 has high water solubility, a dilute sulfuric acid solution is required to dissolve Ti2(SO4)3. TiOSO4 is probably formed at a low calcination temperature and is converted to Ti2(SO4)3 at 400 °C, but the sufficiently low pH of the leaching solution enables both Ti phases to dissolve easily during the water leaching process.

[0207] Example 2: Recovery of Valuable Elements from BF Slag by ABWL (a) Preparation and Characterization of BF Particles The BF slag sample was crushed and ground using a ball mill to obtain BF slag particles. The BF slag particles were then sorted and classified into narrow-width fractions to obtain substantially uniform particle sizes of approximately 200 mesh.

[0208] The particle size distribution of the pulverized BF slag was calculated using a particle size analyzer, and the results are shown in Figure 11. As can be seen in Figure 11, the median particle size is approximately 44.75 μm, and the average particle size is approximately 51.23 μm. 10 It is approximately 11.44 μm, D 90 Its diameter is approximately 98.76 μm.

[0209] We characterized BF slag particles to identify valuable substances in the BF slag, obtained baseline concentrations, calculated extraction efficiency, and elucidated the reaction mechanism, focusing on structural, physical, and morphological changes.

[0210] The elemental composition of BF slag was characterized by ICP-OES and ICP-MS after aqua regia immersion (Figures 12(a) and (b)). ICP-OES was used for the quantification of base metals, while ICP-MS was used for REE analysis because the complexity of the matrix during ICP-OES measurement interfered with the accurate and precise quantification of REE in the BF slag leachate.

[0211] As shown in Figures 12(a) and (b), the main components of the BF slag particles are Ca (approximately 21.7 wt.%), Al (approximately 10.3 wt.%), Mg (approximately 9.4 wt.%), and Fe (approximately 1.3 wt.%). Although silica is known to be the main component of BF slag, it was not measured by ICP-OES because aqua regia immersion could not completely dissolve the silica. However, based on the X-ray fluorescence (XRF) results shown in Figure 12(c), the silica content was determined to be approximately 14.3 wt.%. ICP-MS results (Figure 12(b)) showed that the sample contained Sc, Y, La, Ce, Pr, Nd, Sm, Gd, and Dy at a total REE concentration of approximately 355 mg / kg.

[0212] XRD analysis was performed to study the crystal structure of BF slag particles, and the results are shown in Figure 12(c). As shown in Figure 12(c), the BF slag particles consist of three main oxide phases, Ca2(Mg 0.75 Al 0.25 )(Si 1.75 Al 0.25 O7)(okermanite), CaMgSiO4 (monchelite), and (Ca,Ce)(Ti,Fe 3+ ,Cr,Mg) 21 O 38 (Including loveringite). These results are in good agreement with the ICP-OES results.

[0213] SEM-EDS analysis was performed to examine the physical and morphological properties of the BF slag particles (Figure 12(d)). The SEM results indicate that the BF slag particles have a rough, porous surface.

[0214] Elemental mapping results showed that Ca, Si, and Al were highly correlated with each other, while the degree of correlation of Mg with these elements appeared lower. Fe was distributed almost uniformly across the particle surface at low concentrations, while Mg and Ti were concentrated in specific regions. REE was not detected by XRD and EDS elemental mapping due to its low concentrations.

[0215] (b) Recovery of metals from BF slag particles This example demonstrates the recovery of metal from BF slag particles. The selected BF slag particles from Example 2(a) were dried in a convection oven for more than 24 hours.

[0216] After drying, the selected BF slag particles described in Example 2(a) were mixed with concentrated (96 wt.% to 98 wt.%) H2SO4 and then calcined in a box furnace at a temperature of 200°C to 400°C. The acid-calcined samples were then leached in water at 25°C using a magnetic stirrer for 6 hours. The leached solutions were collected at specified time intervals, diluted, and then characterized by ICP-OES and ICP-MS.

[0217] The BF slag particle samples were characterized using XRD and SEM-EDS before and after acid roasting and water leaching.

[0218] (c) Experimental design and empirical model construction A systematic study was conducted to examine the quantitative effects of six operating factors, namely, acid roasting temperature (X1), acid-to-slag particle mass ratio (X2), acid roasting time (X3), water-to-slag particle density (X4), stirring speed (X5), and water-to-slag mass ratio (X6), and their combinations, on the extraction efficiencies of REE (Sc and Nd) and base metals (Ca, Al, Mg, and Fe).

[0219] For statistical examination, the test factors were coded into low (-1), medium (0), and high (+1) levels to enable a direct comparison of the relative effects of each factor based on the magnitude of the factor model coefficients. Detailed information on the factor levels is shown in Table 6. The upper and lower limits of the factor levels were selected based on Example 1(c) and preliminary experiments.

Table 6

[0220] To construct an empirical extraction model, at a constant leaching time (360 minutes)

Number

Table 7

[0221] The model in the form shown in Equation 18 was then fitted to the extraction efficiency data by mMLSR. A simplified model containing only significant parameters (α=0.05) was evaluated by ANOVA to appropriately estimate the variance of the run. To better illustrate the kinetic effect on extraction efficiency at different temperatures, rate tests were performed at the lower and upper limits of the firing temperature, while keeping other parameters at intermediate factor levels.

number

[0222] A detailed explanation of the statistical analysis is provided in Example 1(c).

[0223] (d) Speed ​​experiment Two rate experiments were conducted, one with X1 at 200°C (-1 level) and the other at 400°C (+1 level), with other factors set to a medium level (X2 = 1.25 g / g). DBFS、 X3=60 minutes, X4=10mL / g ABBFS X5 = 400 rpm, and X6 = 1 g / g DBFS The process was carried out while maintaining the temperature, and the effect of time on extraction efficiency at different firing temperatures was investigated.

[0224] As shown in Figure 13, at level -1 X1, all elements except Nd reached equilibrium within 6 hours. Al and Mg showed fast kinetics, reaching a plateau in 1 hour, while the extraction efficiencies of Fe and Sc increased only slightly. The Nd concentration continued to increase over time, indicating that it had the slowest kinetics of the elements studied. When X1 changed to level +1, the maximum extraction of Sc and Al was obtained within 2 hours of leaching. The extraction kinetics of the other elements were slower, and therefore more than 6 hours were required to reach equilibrium. In general, the extraction efficiencies of base metals were higher than those of REE, and Sc showed higher extraction compared to Nd.

[0225] At a 6-hour leaching time, an increase in X1 was found to have a negative effect on the extraction efficiency of all elements except Al. Specifically, the extraction of Sc, Nd, and Fe was significantly affected by this factor, with changes in extraction efficiency from 52.6%, 32.8%, and 71.1%, respectively, to 35.8%, 22.8%, and 57.2%, respectively.

[0226] The results of this kinetic study showed that leaching time significantly affects the extraction process. For further experiments, the leaching time was set to 360 minutes to ensure that equilibrium was reached in most cases.

[0227] (e) Effects of operational parameters and empirical model building Experiments were conducted to examine the quantitative effects of six operational parameters and the quadratic interactions between at least two of these parameters.

[0228] A total of 19 runs were performed, including 3 center point runs, and leachate samples were collected after 360 minutes of leaching. Figure 14 shows the quantitative effects of principal factors and secondary interactions based on the ordered factor effect coefficients of the empirical model for each element. Only factors with sufficient significance (α=0.05) are shown.

[0229] As shown in Figure 14(a), the extraction efficiency of Sc is positively affected by X2, while X 12 +X 35 (Castration temperature and acid-to-slag mass ratio, calcation time and stirring rate) and X1 had negative effects. The positive effect of X2 is understandable, as the amount of acid controls the phase conversion from water-insoluble oxides and silicates to water-soluble sulfates. The effect of X1 is consistent with the results of rate tests, which have already shown the adverse effect of X1 on the extraction efficiency of most elements except Al. This effect may be due to the decomposition of sulfuric acid at high temperatures. At temperatures raised above 250°C, it was found that liquid sulfuric acid either changes to the gas phase or decomposes into SO3, SO2, H2O, and O2 gases.

[0230] As shown in Figure 14(b), the extraction efficiency of Nd is X2, X4, and two secondary interactions (X 24 +X 36 and X 13 +X 25 ) is positively influenced by X1, X3, X6, and several quadratic interactions (X 26 +X 34 , X 15 +X 23 +X 46 , and X 12 +X 35 ) was negatively affected by this.

[0231] First, at higher water ratios, the solution is unsaturated, and therefore more salt can be dissolved. Longer periods and higher calcination temperatures can lead to more active decomposition and phase transitions of sulfuric acid, thereby hindering extraction. The negative effect of X6 can be explained by the fact that adding water to the slag and acid mixture lowers the boiling point of sulfuric acid, and therefore hinders the reaction between the acid and slag in the acid calcination step. When +1 level X2 and X6 are used, the final concentration of the acid decreases from 98% to 49%, and the boiling point decreases from about 337°C to about 195°C. Thus, even if X1 is at -1 level, the acid evaporates before reaching the set temperature and reacting with the BF slag particles.

[0232] As shown in Figure 14(c), the extraction efficiency of Al depends on only two factors, namely X2 and X4. As shown in Figures 14(d) and (e), the extraction efficiency of Mg and Fe depends on X2, X3, X4, and several secondary interactions (X 26 +X 34 and X 16 +X 45 Positively influenced by (including), while X1, X6, and a small number of quadratic terms (X 24 +X 36 and X 15 +X 23 +X 46was negatively affected by (including). Furthermore, unlike other elements, the dissolution of Mg was improved with the increase of X6. An overview of the operating conditions and the corresponding extraction efficiencies of Sc, Nd, Ca, Al, Mg, and Fe is provided in Table 8 below.

Table 8

[0233] (f) Verification and Optimization An empirical extraction model was constructed for each element using Equation 18 to predict the extraction efficiency under specific conditions. As shown in Figure 15, two verification tests were conducted to evaluate the applicability of the model. The first verification test was at X1 = 400 °C, X2 = 2 g / g DBFS , X3 = 30 minutes, X4 = 16 mL / g ABBFS , X5 = 200 rpm, and X6 = 2 g / g DBFS . The results were consistent with the model predictions, and the model accuracy was 7.7%. The second verification test was conducted under mild conditions (X1 = 200 °C, X2 = 0.5 g / g DBFS , X3 = 30 minutes, X4 = 4 mL / g ABBFS , X5 = 600 rpm, and X6 = 0 g / g DBFS ). Using these conditions, the actual extraction efficiencies of REE and Al were less than 11%, and the model accuracy was determined to be 15.8%.

[0234] Optimization tests were conducted to maximize the economic benefits obtained from all the studied elements. The conditions that satisfy this optimization were X1 = 200 °C, X2 = 2 g / g DBFS , X3 = 90 minutes, X4 = 16 mL / g ABBFS , X5 = 600 rpm, and X6 = 0 g / g DBFS . The predicted benefit from all elements was 264.3 USD / ton DBFS , and the actual benefit was calculated to be 238.1 USD / ton DBFS (accuracy = ±20.4%). The prices of Sc, Nd, Al, Mg, and Fe considered in the optimization calculation are shown in Table 9.

Table 9

[0235] An optimization test was conducted to maximize the economic benefits derived from REE.

[0236] The conditions that satisfy this optimization are X1 = 200°C and X2 = 2g / g DBFS , X3=30 minutes, X4=16mL / g ABBFS X5 = 600 rpm, and X6 = 0 g / g DBFS As shown in Figure 15, the actual extraction efficiency was similar to that predicted with high accuracy (±12.8%). The predicted and actual profits considering REE were 69.0 and 71.1 USD / ton, respectively. DBFS It was calculated to be this.

[0237] The overall validation and optimization results demonstrate that the empirical model can successfully predict optimal conditions and extraction efficiency.

[0238] Example 3: Optimized parameters for the recovery of individual target metals by ABWL (a) Recovery of valuable elements from EAF slag Table 10 below shows the parameters optimized for the recovery of Nb, Ti, Fe, Ca, Mn, Mg, and Al from EAF slag, as determined in Example 1. [Table 10]

[0239] (b) Recovery of valuable elements from BF slag The parameters optimized for the recovery of Sc, Nd, Ca, Al, Mg, and Fe from BF slag, as determined in Example 2, are shown in Table 11 below. [Table 11]

[0240] Example 4: Recovery of valuable elements from EAF slag by carbonothermal reduction followed by ABWL (a) Preparation and characterization of EAF slag particles First, the EAF was crushed and ground using a jaw crusher and disc to obtain a substantially uniform particle size of approximately 200 mesh (approximately 74 μm). The ground slag sample was dried in a convection oven at a temperature of 50°C for more than 24 hours to remove moisture content.

[0241] Next, the sample was mixed with lignite and a flux (SiO2 and / or Al2O3), and the mixture was then pelletized using an air / hydraulic press at a pressure of approximately 250 MPa for 3 minutes (pellet diameter = 28.6 mm). The pellets were then placed in a graphite crucible and smelted in a box furnace under an argon atmosphere at different set temperatures (1500°C or 1600°C, heating rate 180°C / hour, cooling rate 180°C / hour, holding time 1.5 hours). After smelting, the metallic Fe phase and slag phase were separated manually using a Dremel® and a hammer, and the slag phase was pulverized using a mixer mill. To reduce the Fe content in the slag phase, the pulverized slag sample was magnetically separated using a Davis® tube. The Fe-depleted slag samples obtained in this manner were then dried in a convection oven and, for characterization, were immersed in an aqueous solution using a microwave immersion apparatus (heating time 40 minutes, holding time 30 minutes, set temperature 220°C).

[0242] Dried Fe-depleted slag samples were then mixed with concentrated sulfuric acid. The mixture was placed in a porcelain crucible and calcined in a muffle furnace. The acid-calcined samples were then leached with water at ambient temperature and pressure using a magnetic stirrer. The resulting leaching solution was then filtered and diluted using a Hamilton® Microlab 600® dual dilution and dispensing system for characterization. Specific solid samples before and after each process were also prepared for post-characterization.

[0243] (b) Experimental design and empirical model building For the carbonothermal reduction process, an experimental matrix was designed using a partial factorial design method, an empirical model was constructed, and the process was optimized as a result. A systematic study was conducted to examine the effects of four operational parameters—smelting temperature (X1), carbon-to-slag mass ratio (X2), flux-to-slag mass ratio (X3), flux type (X4), and the secondary interaction of these factors—on the composition of the slag phase. A low smelting temperature (X1) was selected based on preliminary test results to ensure complete melting of the pellets and clear phase separation between the metallic Fe phase and the slag phase. Stoichiometric calculations showed that at least 0.06 g of carbon per 1 g of EAF slag is required to reduce all iron in the slag in the form of FeO; therefore, a low carbon-to-slag mass ratio (X2) of 0.06 g / g was chosen. EAFS The setting was determined to be silica or alumina due to the high calcium content in the EAF slag.

[0244] For the ABWL process, the effects of four operational parameters—calcination temperature (X1), acid-to-slag mass ratio (X2), calcination time (X3), and water-to-slag ratio (X4)—on the extraction of target elements (Ti, Fe, Ca, Mn, Mg, Al, Cr, and Sr) were investigated using response surface spectroscopy.

[0245] For this process, factor-level boundaries were selected based on preliminary test results and the operating range of similar processes in the literature, within an operating range where the system's response could be relatively linear (Anawati et al., Waste Management 95 (2019): 549-559; Demol et al., Hydrometallurgy 179 (2018): 254-267; Kim et al., Hydrometallurgy 191 (2020): 105-203; Meshram et al., Journal of Cleaner Production 157 (2017): 322-332; Sadri et al., International Journal of Mineral Processing 159 (2017): 7-15; Safarzadeh et al., Mining, Metallurgy & Exploration 29, no. 2 (2012): 97-102; and Safarzadeh et al., International Journal of Mineral Processing 124 (2013): 128-131).

[0246] Based on partial factorial design and response surface methodology, two experimental matrices were designed for the carbonothermal reduction process and the ABWL process, respectively. For the carbonothermal reduction process, a series of

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[0247] (c) Mineralogical, morphological, and compositional property evaluation The mineralogical properties of the solid samples were examined by XRD. Morphological and compositional analyses were performed by SEM-EDS. An electron beam microanalyzer was used for the compositional property evaluation of the solid samples. The elemental concentrations in the liquid samples were measured by ICP-OES. The silica content of the solid samples was quantified by XRF. The carbon content of the metallic Fe phase was analyzed using LECO (CS 444 carbon-sulfur analyzer).

[0248] (d) EAF slag property evaluation The chemical composition of EAF slag was investigated by ICP-OES and XRF analysis after aqua regia immersion (Table 12). As shown in Table 12, the main component of EAF slag is Fe, which accounts for 24.3 wt.% of the composition. The high Fe content confirmed that the EAF slag sample is suitable for undergoing a carbonothermal reduction process. As shown in Table 12, EAF slag also contains Ca (17.3 wt.%), Si (7.5 wt.%) as measured by XRF), Mg (5.5 wt.%), Mn (5.5 wt.%), Al (2.5 wt.%), and Cr (1.9 wt.%). EAF also contains trace amounts of Ti (1862 mg / kg), Na (1184 mg / kg), Nb (489 mg / kg), Sr (341 mg / kg), Cu (303 mg / kg), and Ni (300 mg / kg). [Table 12]

[0249] (e) Effects of operational parameters and empirical model building Carbon thermal reduction was performed to selectively separate Fe, Nb, Cr, Mn, and partially Ti, reporting them in the metallic phase while reporting other elements in the slag phase. Based in part on factorial planning,

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[0250] Figure 16 shows ordered factor effect coefficients with sufficient significance (α=0.05) for the concentrations of each element in the slag phase. It should be noted that the goal of this process is to minimize the concentrations of the five elements in the slag phase so that they are expected to be reported in the metallic phase.

[0251] As shown in Figure 16(a), the Fe concentration in the slag phase was negatively affected by three main factors: X1, X2, and X3. In other words, an increase in these factors resulted in the desired decrease in Fe concentration in the slag phase. The negative effect of X1 is understandable, as a higher X1 reduces the viscosity of the slag and, as a result, improves the separation of the metal phase from the slag phase. The inverse correlation between X1 and slag viscosity is shown in Equation 22.

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[0252] It was observed that the Fe concentration in the slag phase increased with decreasing X2. Low levels of X2 (0.06 g / g) EAFS This means that this level was selected based on the stoichiometric value of carbon needed to reduce all the Fe in the slag, but it was not enough to reduce all the Fe. This phenomenon can be explained by the formation of Mn and Cr carbides that occurs during the smelting process.

[0253] EPMA elemental mapping of the metallic phase was obtained after the carbon-thermal reduction process (run 10), and the results are shown in Figure 17. The results showed fairly strong correlations between regions rich in C, Mn, and Cr, where the Fe concentration was relatively low. This indicates that Fe-Mn-C, Fe-Cr-C, and Fe-Mn-Cr-C alloy formation reactions occurred, consuming a considerable amount of carbon present in the mixture. Based on the chemical composition of the EAF slag, it was calculated that at low levels of X2 (0.06 g of carbon per gram of EAF slag), the FeO phase present in the EAF slag could not be completely reduced to metallic Fe if more than 30 wt.% of Mn and Cr were involved in the carbide formation reaction. However, this reaction did not affect the FeO reduction rate when X2 was at medium levels (0.09 g of carbon per gram of EAF slag) and high levels (0.12 g of carbon per gram of EAF slag).

[0254] The negative effect of X3 can be explained by the basicity of the slag. The basicity (R) of the four variables was calculated based on Equation 23:

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[0255] If no flux is added, the slag phase obtained after the carbonothermal reduction process has substantially excessive basicity (R=2.584) due to the high basic oxide content in the EAF slag. This substantially excessive basicity can increase the melting temperature of the EAF slag, which results in solid precipitates, namely CaO and MgO, as well as the high apparent viscosity of the EAF slag. Thus, metallic iron droplets can be more easily trapped in the slag. The flux used in this process is either acidic (SiO2) or amphoteric (Al2O3). Given the basic conditions in the system and the low alumina content (<15 wt%), Al2O3 is likely to act as an acidic oxide. Adding these acidic fluxes can, for example, lower the basicity and liquidus point of the EAF slag, resulting in a clear separation between the metallic phase and the slag phase. Thus, the Fe content in the slag phase decreases in the presence of SiO2 or Al2O3. To verify these explanations, 0.09 g / g EAFS Using FactSage software, the carbon-to-slag mass ratio (0 level) was simulated for the carbon-to-slag reduction of EAF slag samples with and without a flux (SiO2 or Al2O3). The results confirmed that the addition of either SiO2 or Al2O3 as a flux significantly lowered the liquidus point of the slag.

[0256] The concentrations of other elements in the slag phase showed a similar trend compared to that of Fe. However, the type of flux (X4) showed a further positive effect on the concentrations of Nb and Mn. The concentrations of Nb and Mn in the slag phase decreased when SiO2 was used as the flux instead of Al2O3 (Figures 16(b) and (d)). This may be due to the fact that SiO2 is a stronger acidic oxide than Al2O3. The presence of SiO2 contributes to lowering the liquidus level of the EAF slag and improving phase separation, while Al2O3 has the same effect, but to a lesser degree. Furthermore, it was found that X2, unlike the other elements, had a positive effect on the concentration of Mn in the slag phase. This can be explained by the desulfurization of the metal by Mn. As described in Example 4(a), lignite was used as a reducing agent, and this type of coal contains about 0.4–1.0 wt.% sulfur, making it the main sulfur source. Substantially high sulfur content can make metal products brittle, poorly weldable, and susceptible to corrosion; therefore, it is desirable to maintain the sulfur content below 0.015 wt.% (Schrama et al., Ironmaking & Steelmaking 44, no. 5 (2017): 333-343). MnO, along with CaO and MgO, is one of the main desulfurizing agents, used to transfer sulfur from metal to slag, following the reaction shown below (Equation 24): FeS (metal) +Mn (metal) →Fe (metal) +MnS (slag) [Formula 24]

[0257] When a larger amount of carbon (higher levels of X2) is introduced into the system, more sulfur is introduced and reacts with the Fe metal, resulting in the production of more FeS. According to Le Chatelier's principle, desulfurization occurs actively through the formation of MnS in the slag phase.

[0258] The metallic phase obtained after carbonothermal reduction was characterized using XRD and EPMA analysis. The X-ray diffractogram showed that the metallic phase consisted of Fe, Fe3C, and C, confirming a clear separation between the metallic phase and the slag phase. The presence of niobium phosphide was observed by EPMA.

[0259] (f) Optimization and verification of the carbon thermal reduction process Based on experimental results and factor effect coefficients, an empirical model was constructed to predict the concentrations of Fe, Nb, Cr, Mn, and Ti in the slag phase. The empirical model for Fe concentration is shown in Equation 25:

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[0260] Two validation tests were conducted to evaluate the predictive power of the model. As shown in Table 14, the results demonstrated that the model had high model accuracy in predicting the concentrations of five elements (Fe, Nb, Cr, Mn, and Ti) in the slag phase (mean absolute relative deviation (AARD) of 15.6% and 12.1% for each test set). [Table 14]

[0261] The carbon-thermal reduction process was then optimized to minimize the Fe concentration in the slag phase. The optimized conditions were a high smelting temperature (X1, 1600°C) and a carbon-to-slag mass ratio (X2, 0.12 g / g). EAFS ), as well as low levels of flux-to-slag mass ratio (X3, 0g / g EAFS The materials used were the metal and the type of flux (X4, silica). The minimum amount of Fe in the slag phase at a concentration of 1620 mg / kg was obtained after the smelting process under optimal conditions, along with metallic iron rich in Mn, Cr, and Nb (approximately 70 wt.% Fe, approximately 4.14 wt.% C, approximately 17 wt.% Mn, approximately 7 wt.% Cr, and approximately 0.2 wt.% Nb) (Table 14). Table 15 shows the composition of the raw material EAF slag, as well as the composition of the metallic phase and slag phase obtained after the carbon-thermal reduction process under optimal conditions. [Table 15]

[0262] (g) Results and characterization of water leaching from acid calcination of Fe-depleted slag The slag generated after the carbonothermal reduction process under optimal conditions was thoroughly characterized using ICP-OES, XRD, EPMA, and SEM-EDS after aqua regia immersion (Figure 18). Most Fe, Nb, Cr, Mn, and partially Ti were reported in the metallic phase, while the concentrations of these elements in the slag phase were minimal.

[0263] Figure 18(a) shows the elemental composition of the Fe-depleted slag. The slag sample mainly consisted of Ca (30.2 wt.%), Mg (11.7 wt.%), and Al (4.9 wt.%), with trace amounts of Mn (6803 mg / kg), Ti (2993 mg / kg), Fe (1620 mg / kg), Sr (628 mg / kg), and Cr (175 mg / kg). Based on the XRD results shown in Figure 18(b), the slag consisted of four phases: lanite (Ca2SiO4), magnesium iron oxide (Mg 0.91 Fe 0.09 It consists of O), calcite (CaCO3), and gehlenite (Ca2Al2SiO7). EPMA phase mapping results further confirm that the slag contains these four phases along with small amounts of chromite contributing to trace amounts of Fe and Cr (Figure 18(c)). EDS elemental mapping in Figure 18(d) shows that Ca and Si are highly related to each other, confirming the presence of Ca2SiO4. Aluminum is also present in some cases within the Ca and Si-rich regions, indicating Ca2Al2SiO7. Mg matches only with oxygen, and Mg 0.91 Fe 0.09 The presence of O is confirmed. The SEM image shows that Ca2Al2SiO7 has a smooth surface and a sharp shape, while Ca2SiO4 and Mg 0.91 Fe 0.09 O has a layered structure (Figure 18(d)).

[0264] (h) Effects of operational parameters and empirical model building For the ABWL process, a circumscribed center composite experimental matrix was designed using the response surface method, and a second-order empirical extraction model for the target element was constructed. Detailed experimental conditions, as well as the corresponding extraction efficiencies for Ti, Fe, Ca, Mn, Mg, Al, Cr, and Sr, are shown in Table 16. For the extraction efficiency of each element, ordered factor effect coefficients with sufficient significance (α=0.05) are shown in Figure 19. [Table 16] TIFF0007880397000039.tif16162

[0265] As shown in Figure 19, the acid-to-slag mass ratio (X2) showed a large positive effect on the extraction of all elements. This may suggest that H2SO4 is the limiting reagent for the immersion reaction. At an acid-to-slag ratio of 0 (1.5 g / g), it was calculated that there was an equivalent amount of H2SO4 in the system that could completely react with the slag sample, and therefore, the immersion reaction did not occur sufficiently at lower acid ratios (0.5 g / g and 1 g / g).

[0266] Unlike other elements, Ca and Sr were also positively affected by the water-to-slag ratio (X4) (Figures 19(c) and 19(h)). This may be due to the low water solubility of CaSO4 and SrSO4. The solubility of CaSO4 and SrSO4 in water is 2.00 g / L and 0.117 g / L, respectively, which is less than 1 / 100th of the solubility of common salts. Therefore, a higher water-to-slag ratio was required to increase the extraction of Ca and Sr.

[0267] The calcination temperature (X1) showed a negative effect on the extraction of all elements of interest. This effect can be explained, for example, by the phase transition and decomposition of H2SO4 at high calcination temperatures. Above 250°C, liquid H2SO4 begins to evaporate and decompose into SO2, SO3, H2O, and O2 gases, and therefore liquid H2SO4 cannot exist at temperatures above 350°C. This phenomenon results in a substantial decrease in the amount of liquid H2SO4, which is highly reactive with slag particles. Calcination temperature is potentially the main contributing factor to mass loss, and the adverse effect of this factor is confirmed. Higher levels of acid-to-slag mass ratio (X2), calcination time (X3), and calcination temperature × acid-to-slag mass ratio (X 12 ) and firing temperature × firing time (X 13 The secondary interaction of ) also revealed that it accelerates the mass reduction of the acid and slag mixture. This is because X3, X1 together with X1. 12 , and X 13 However, in most cases, this explains why a greater mass loss occurs at higher levels of X2, while a higher absolute amount of available H2SO4 molecules causes a positive effect of this factor on the extraction.

[0268] Figure 20 shows a 2D contour plot visualizing the interaction effect of two principal factors on the average extraction efficiency of all elements except Ca. The extraction efficiency of Ca was not considered because this process was designed to recover Ca as gypsum (CaSO4·2H2O) in the residue. As can be seen, the acid-to-slag mass ratio (X²) plays a significant role in increasing the overall extraction efficiency, while the other factors do not have a major impact on extraction.

[0269] (i) Optimization of the acid calcination water leaching process Based on the factor effect coefficients obtained after the design-of-experiment tests, the ABWL process was optimized to maximize the extraction of Ti, Fe, Mn, Mg, Al, Cr, and Sr. Calcium extraction efficiency was not considered in the process optimization because Ca is recovered as gypsum (CaSO4·2H2O) in the residue.

[0270] The response surface method allows for the exploration of a wider range of response surfaces, as long as the Euclidean distance remains constant. Therefore, unlike some factorial design methods, the optimized factor levels are not limited to -1 and 1, but can be any number. However, it should be noted that while the predicted extraction efficiency may exceed 100% under certain conditions, the actual efficiency cannot exceed 100%, and therefore, extraction efficiency ≤ 100% was added as a constraint for process optimization. Two optimized conditions were then determined with and without the constraint of factor levels limited to -1 and 1. Figure 21 shows the optimization results (i.e., extraction efficiency) of the acid calcined water leaching process. The results showed that extraction efficiencies of over 80% were achieved for almost all elements except Ca and Sr, regardless of the factor level constraint. The low efficiencies for Ca and Sr may be due to their low water solubility.

Claims

1. A method for recovering at least one target metal from iron or steel slag (hereinafter referred to as "steel slag, etc.") from an electric arc furnace or a basic oxygen furnace (BOF), comprising the following steps: (a) A step a in which a mixture is produced by mixing particles such as steel slag with at least one reducing agent having a mass ratio of the reducing agent to the particles such as steel slag in the range of 0.06 to 0.12 and at least one flux having a mass ratio of the flux to the particles such as steel slag in the range of 0 to 0.

1. The flux is selected from the group consisting of silica, alumina, and combinations of silica and alumina, step a, (b) The mixture is smelted at a temperature of 1300°C to 1800°C to form a metallic phase and a slag phase containing iron as the first target metal, step b, (c) Separating the metal phase containing the first target metal from the slag phase to produce the metal phase containing the target metal and iron-depleted slag, step c A method that includes

2. The method according to claim 1, wherein the flux is alumina, or a combination of alumina and silica.

3. The reducing agent includes a carbon source, or The aforementioned particles, such as steel slag, have a size of less than 200 mesh. The method according to claim 1 or 2.

4. below: Prior to step a, a step of crushing the particles such as steel slag; A step of classifying and separating the particles such as steel slag into fractions according to their size; Prior to step a, a step of drying the particles such as steel slag; Prior to step b, a step of pelletizing the mixture; A step of crushing the slag phase to obtain slag particles; and The step of reducing the iron content in the slag phase to produce iron-depleted slag. The method according to any one of claims 1 to 3, further comprising at least one step selected from the group consisting of the following.

5. The method according to any one of claims 1 to 4, wherein the separation in step c is carried out by a mechanical separation method.

6. The method according to any one of claims 1 to 5, wherein the smelting in step b is carried out at a temperature of 1500°C to 1600°C.

7. The method according to any one of claims 1 to 6, wherein the particles such as steel slag and the reducing agent undergo a redox reaction that releases chemical energy.

8. The method according to any one of claims 1 to 7, wherein step a includes mixing at least one flux having a mass ratio of the flux to particles such as steel slag of 0.

05.

9. The method according to any one of claims 1 to 7, wherein step a includes mixing at least one flux having a mass ratio of the flux to particles such as steel slag of 0.

1.

10. The method according to any one of claims 1 to 9, wherein the reducing agent is a carbon source selected from the group consisting of metallurgical coal, charcoal, petroleum coke, petroleum coke, natural gas, and at least two combinations thereof.

11. The method according to any one of claims 1 to 10, wherein the mass ratio of the reducing agent to the particles such as steel slag is 0.06, 0.09, 0.10, or 0.

12.

12. The method according to any one of claims 1 to 11, comprising the step of drying the particles such as steel slag before mixing, wherein the drying is carried out at a temperature of at least 50°C and for a period of at least 24 hours.

13. The method according to any one of claims 1 to 12, wherein step b is performed in a temperature range of 1300°C to 1700°C, 1300°C to 1600°C, 1400°C to 1800°C, 1400°C to 1700°C, 1400°C to 1600°C, 1500°C to 1800°C, 1500°C to 1700°C, or 1500°C to 1600°C.

14. The process further includes a step of subjecting the aforementioned iron-depleted slag to a high-temperature wet smelting process to recover a second target metal, and The step of subjecting the aforementioned iron-depleted slag to a high-temperature wet smelting process to recover a second target metal is as follows: A step of mixing the iron-depleted slag particles and an acid such that the mass ratio of the acid to the iron-depleted slag particles is in the range of 0.5 to 5 to produce a further mixture, wherein the acid is selected from the group consisting of sulfuric acid, hydrochloric acid, and a mixture of at least two of these. The further mixture is calcined at a temperature of 100°C to 600°C to remove excess water and the acid, and a dry mixture containing pyrolysis gas and at least one soluble metal salt is produced. The steps include adding water to leach the dry mixture to a density in the range of 50 g / L to 250 g / L, thereby producing a mixture containing an aqueous leachate rich in the second target metal and a solid residue, The steps include separating the aqueous leachate rich in the second target metal from the solid residue, and The method according to any one of claims 1 to 13, comprising the above.

15. The method according to claim 14, wherein the second target metal is selected from the group consisting of titanium, niobium, manganese, chromium, scandium, neodymium, yttrium, lanthanum, cerium, samarium, gadolinium, dysprosium, praseodymium, europium, terbium, erbium, calcium, magnesium, aluminum, copper, silicon, ruthenium, rhodium, palladium, osmium, iridium, and platinum.

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