Method for oxygen transfer refining of liquid iron and top-blown lance

Abstract

A molten iron oxygen-supply refining method in which, during at least a part of the oxygen-supply refining, an oxygen-containing gas is jetted from a jetting nozzle of an outer shell of a through-top blowing lance, while a control gas is jetted from a jetting outlet (3) toward the inside of the jetting nozzle, and the oxygen-containing gas as a main supply gas is supplied from the inlet side of the jetting nozzle, the jetting outlet (3) is provided to the nozzle side surface of a position (1) where the cross-sectional area of the nozzle is the smallest in the nozzle axial direction or a position near the position, and is configured in such a manner that, when an arbitrary plane passing through the center axis of the nozzle is divided into two spaces, at least a part of the jetting outlet exists in both spaces.

Description

[0001] This application is a divisional application of patent application 201880080103.X (application date: November 8, 2018, invention title: oxygen delivery refining method for molten iron and top blowing lance). Technical Field

[0002] The present invention relates to an oxygen-feeding refining method for molten iron by blowing oxygen-containing gas from a top-blown lance into molten iron loaded into a reaction vessel, and to a top-blown lance used for the oxygen-feeding refining. Background Technology

[0003] In the oxidative refining of molten iron, from the perspective of improving reaction efficiency, there has been a continuous search for a practical oxygen delivery method that can simultaneously control the jet velocity and gas flow rate of oxygen-containing gas injected from the top-blown lance on the surface of the molten iron bath.

[0004] For example, in the decarburization and refining of molten iron using a converter, the top-blown oxygen flow rate per unit time is sometimes increased to improve converter productivity. However, if the flow velocity of the jet from the molten iron bath increases in this case, the amount of iron particles scattered outside the furnace in the form of furnace dust and other substances, as well as those adhering to and depositing on the furnace walls and near the furnace opening, increases. If this amount increases, it leads to increased costs and reduced converter operating rate due to a decrease in the yield of finished iron. Therefore, a method for supplying oxygen that can achieve high flow rate and low velocity is sought.

[0005] On the other hand, when the carbon concentration in the molten iron is low at the end of the blowing process, the top-blown oxygen flow rate is generally reduced to prevent excessive iron oxidation and loss. In this case, if the flow rate of the jet from the molten iron bath is too low, the stirring of the molten iron below the ignition point is weak, leading to excessive iron oxidation. Therefore, an oxygen delivery method is sought that can operate at a low flow rate when the oxygen flow rate is high, and at a high flow rate when the oxygen flow rate is low.

[0006] Generally, adjusting the flow velocity on the bath surface independently of oxygen flow rate adjustment has been achieved by adjusting the lance height. However, if the lance height is too low, molten iron splashing out will cause melting and damage, significantly reducing the lance's lifespan. Conversely, if the lance height is too high, the furnace gas temperature will rise due to increased secondary combustion rate and decreased secondary combustion thermal efficiency, potentially reducing the refractory's lifespan. Therefore, the range of flow velocity adjustment based on lance height is limited. Thus, there is a need for an oxygen delivery nozzle capable of adjusting the injection velocity independently of oxygen flow rate.

[0007] However, the gas velocity at the nozzle outlet generally exhibits the following property: if the nozzle shape is determined, the gas velocity is uniquely determined relative to the gas flow rate; the velocity increases at high flow rates and decreases at low flow rates. In particular, if the nozzle diameter is increased to achieve low dynamic pressure at high gas flow rates, there is a problem that the velocity drops too low when the gas flow rate is reduced. Therefore, a technique has been investigated that, by controlling the nozzle shape during blowing, it is possible to simultaneously achieve blowing conditions where the dynamic pressure does not become excessively high at high oxygen flow rates and blowing conditions where the dynamic pressure does not become excessively low at low oxygen flow rates. For example, Patent Document 1 discloses a technique for controlling the nozzle shape in a vacuum degassing tank using a top-blown spray gun that mechanically changes the nozzle shape.

[0008] Furthermore, Patent Document 2 discloses an operating method using a Laval nozzle, in which a gas outlet hole is provided on the inner surface of the expanded portion at the end of the Laval nozzle, and gas is blown in through this outlet hole according to the mainstream oxygen flow rate. Laval nozzles are widely used in converter refining because they can efficiently convert gas pressure into kinetic energy, ensuring sufficient gas velocity at the molten iron bath surface even when the converter lance height is increased. In a Laval nozzle, the pressure ratio between the nozzle inlet and outlet is determined by the ratio of the cross-sectional area of ​​the nozzle outlet to the throat (the area of ​​the cross-section perpendicular to the central axis of the nozzle) (opening ratio), which allows for appropriate expansion at the expanded portion of the nozzle to minimize energy loss. Since the furnace pressure at the nozzle outlet is typically atmospheric pressure, the gas pressure at the nozzle inlet (appropriate expansion pressure) and the corresponding gas flow rate (appropriate expansion flow rate) are uniquely determined relative to the nozzle shape. However, if the gas flow rate is reduced to below the appropriate expansion flow rate, the gas pressure at the nozzle inlet will be lower than the appropriate expansion pressure (appropriate expansion pressure), resulting in an over-expansion state that generates shock waves within the nozzle. Conversely, if the gas flow rate is increased to above the appropriate expansion flow rate, it will result in an under-expansion state that generates shock waves below the nozzle outlet, leading to energy loss and a lower gas velocity compared to the case where the nozzle shape achieves appropriate expansion at various gas pressures.

[0009] In the method of Patent Document 2, a small amount of gas is blown into a gas outlet hole located on the inner surface of the end enlargement of the Laval nozzle at a gas flow rate lower than the appropriate expansion flow rate. This causes the airflow forming the boundary layer along the nozzle side of the end enlargement to be pushed inward and stripped away. This suppresses the expansion of the mainstream gas, mitigates over-expansion, and inhibits the decrease in gas velocity when the gas flow rate decreases.

[0010] In addition, as a method for controlling the gas jet flow by blowing gas separately into the nozzle from the mainstream, Patent Document 3 discloses a method for controlling the jet direction of the mainstream gas jet by blowing working gas into the throat of the Laval nozzle in the top-blowing spray gun of the RH degassing equipment.

[0011] Existing technical documents

[0012] Patent documents

[0013] Patent Document 1: Japanese Patent Application Publication No. 8-260029

[0014] Patent Document 2: Japanese Patent Application Publication No. 2000-234116

[0015] Patent Document 3: Japanese Patent Application Publication No. 2004-156083 Summary of the Invention

[0016] Patent Document 1, which describes a method for mechanically changing the nozzle shape, is not only impractical in high-temperature atmospheres where furnace dust is generated due to its mechanically movable parts, but also has the problem of being difficult to apply to spray guns with many nozzle orifices. Furthermore, when the cross-sectional area is reduced using a movable part on the inner surface of the nozzle, a height difference is created in the stepped portion, and the influence of the shape of this height difference on the gas flow rate is unclear.

[0017] In addition, the method in Patent Document 2 aims to separate the boundary layer of the airflow from the nozzle wall by expanding the end of the Laval nozzle, thereby mitigating the excessive expansion state at low gas flow rates. However, it has the following problem: under insufficient expansion conditions where the gas supply pressure is higher than the appropriate expansion pressure determined by the nozzle opening ratio, the flow rate cannot be effectively increased.

[0018] In particular, to improve the productivity of oxygen refining processes such as converters, it is necessary to increase the oxygen flow rate. Sometimes, the nozzle cross-sectional area at the throat is enlarged to suppress gas velocity under high gas flow conditions. However, since it is necessary to ensure an appropriate flow path cross-sectional area for cooling water to cool the nozzle tip, the nozzle outlet cross-sectional area is constrained, and therefore the nozzle opening ratio cannot necessarily be freely set. In this case, there is a tendency for the nozzle opening ratio and the appropriate expansion pressure determined therefrom to decrease, resulting in insufficient expansion conditions even under low gas flow conditions. However, the method in Patent Document 2 cannot effectively increase the gas velocity under such conditions.

[0019] It should be noted that the method in Patent Document 3 has the following problem: even if the direction of gas ejection can be controlled, the gas flow rate cannot be effectively controlled.

[0020] The object of the present invention is to provide a top-blowing oxygen delivery method with a large variable range of gas flow rate, which can effectively increase the gas flow rate even under insufficient expansion conditions without the use of a mechanically movable part in the spray gun nozzle, and the top-blowing spray gun used therein.

[0021] In order to solve the above-mentioned problems, the inventors conducted in-depth research on a method of controlling the gas flow rate without relying on the gas flow rate by changing the method of introducing gas into the nozzle by not setting a mechanical movable part in the top-blown gas injection nozzle. As a result, the oxygen delivery refining method and the top-blown spray gun for the oxygen delivery refining of the present invention were completed.

[0022] That is, the present invention is an oxygen-feeding refining method for molten iron, characterized in that oxygen-containing gas is blown from a top-blown lance into the molten iron loaded into a reaction vessel to perform oxygen-feeding refining. During at least a portion of the oxygen-feeding refining, in the oxygen-containing gas injection nozzle penetrating the outer shell of the top-blown lance, a control gas is ejected from the outlet toward the injection nozzle, while oxygen-containing gas as the main supply gas is supplied from the inlet side of the injection nozzle. The outlet is located on the nozzle side at or near the location where the cross-sectional area of ​​the nozzle is the smallest in the nozzle axial direction, and is arranged such that at least a portion of the outlet exists in both spaces when the nozzle is divided in two by any plane passing through the central axis of the nozzle. Furthermore, as a preferred example, the location near the location where the cross-sectional area of ​​the nozzle is the smallest in the nozzle axial direction is a location where the cross-sectional area of ​​the nozzle is 1.1 times or less of the smallest cross-sectional area in the nozzle axial direction.

[0023] It should be noted that, throughout this specification, the "cross-sectional area" of the nozzle refers to the area inside the nozzle perpendicular to the central axis. Therefore, in this invention, "a portion of less than 1.1 times the minimum cross-sectional area" refers to a portion whose cross-sectional area exceeds 1.0 times but is less than 1.1 times the minimum cross-sectional area.

[0024] It should be noted that, in the oxygen refining method for molten iron of the present invention configured as described above, a more preferred solution is as follows:

[0025] (1) As the injection nozzle, a straight pipe nozzle with a straight pipe section having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, or a Laval nozzle with an end enlargement in the throat having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, is used.

[0026] (2) Ensure that the pressure of the main supply gas at the inlet side of the above-mentioned injection nozzle is greater than the appropriate expansion pressure Po that satisfies the following formula (1):

[0027] Ae / At = (5 5／2 / 6 3 )×(Pe / Po) －5／7 ×[1-(Pe / Po] 2／7 ] －1／2 ···(1)

[0028] Here, At: minimum cross-sectional area of ​​the injection nozzle (mm²)2 Ae: Exit cross-sectional area of ​​the injection nozzle (mm²) 2 Pe: Atmospheric pressure at the nozzle outlet (kPa), Po: Appropriate expansion pressure of the nozzle (kPa).

[0029] (3) The above-mentioned nozzle outlets are arranged circumferentially on the side of the above-mentioned spray nozzle in multiple directions, and the product of the diameter of the inlet hole for introducing the control gas into the above-mentioned nozzle outlets and the number n of the above-mentioned nozzle outlets of each above-mentioned spray nozzle is at least 0.4 times the nozzle inner diameter of the part with the smallest cross-sectional area of ​​the above-mentioned spray nozzle.

[0030] (4) The above-mentioned spray outlet is provided in a slit shape along the entire circumference of the side of the spray nozzle, and the axial length of the spray nozzle at the spray outlet is less than 0.25 times the inner diameter of the nozzle at the part with the smallest cross-sectional area of ​​the spray nozzle.

[0031] (5) During at least a portion of the oxygen refining process, the flow rate of the control gas ejected toward the injection nozzle is at least 5% of the combined flow rate of the control gas and the main supply gas supplied to the injection nozzle.

[0032] (6) Adjust the supply rate of the control gas according to the supply rate of the oxygen-containing gas blown from the top-blown lance to the molten iron.

[0033] (7) As the oxygen refining of the molten iron proceeds, the supply rate of the control gas is changed.

[0034] (8) Based on the silicon concentration of the molten iron before the start of oxygen refining, change the supply rate of the control gas.

[0035] (9) At the end of the oxygen refining process, after 85% of the total oxygen content in the oxygen-containing gas supplied in the above-mentioned oxygen refining process has been reached, the control gas is injected into the above-mentioned injection nozzle while the oxygen-containing gas, which is the main supply gas, is supplied.

[0036] (10) For molten iron with a silicon concentration of 0.40% or more before the start of oxygen refining, in the initial stage of oxygen refining before supplying 20% ​​of the total oxygen content in the oxygen-containing gas supplied in the oxygen refining process, the oxygen-containing gas, which is the main supply gas, is supplied while the control gas is sprayed out of the injection nozzle.

[0037] Furthermore, the present invention is a top-blown spray gun, characterized in that, for blowing oxygen-containing gas into molten iron contained in a reaction vessel, the oxygen-containing gas injection nozzle, which penetrates the outer shell of the top-blown spray gun, has an outlet for ejecting control gas toward the injection nozzle. This outlet is disposed on the nozzle side where the nozzle's cross-sectional area is at or near its minimum cross-sectional area in the nozzle axial direction, and is arranged such that at least a portion of the outlet exists in both spaces when bisected by any plane passing through the nozzle's central axis. Circumferentially arranged on the nozzle side, multiple outlets for introducing control gas into multiple outlets are interconnected within the top-blown spray gun. Furthermore, as a preferred example, the area near the nozzle's minimum cross-sectional area in the nozzle axial direction is a portion where the nozzle's cross-sectional area is 1.1 times or less of the minimum cross-sectional area in the nozzle axial direction.

[0038] It should be noted that, in the top-blowing spray gun of the present invention configured as described above, a more preferred solution is as follows:

[0039] (1) The above-mentioned nozzle outlets are arranged circumferentially on the side of the above-mentioned injection nozzle in multiple directions, and the product of the inner diameter of the nozzle for discharging the control gas that communicates with the above-mentioned nozzle outlets and the number n of the above-mentioned nozzle outlets of each above-mentioned injection nozzle is at least 0.4 times the nozzle inner diameter corresponding to the minimum cross-sectional area of ​​the above-mentioned injection nozzle.

[0040] (2) As a jet nozzle, a straight pipe nozzle with a straight pipe section having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, or a Laval nozzle with an end enlargement in the throat having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, is used.

[0041] Furthermore, the present invention is a top-blown spray gun, characterized in that it is used to blow oxygen-containing gas into molten iron contained in a reaction vessel. The oxygen-containing gas injection nozzle, which penetrates the housing of the top-blown spray gun, has an outlet for ejecting control gas into the injection nozzle. This outlet is provided circumferentially on the nozzle side at or near a location where the cross-sectional area in the nozzle axial direction is the smallest, and is provided in a slit-like manner throughout the entire circumference. Additionally, as a preferred example, the location near the location where the nozzle's cross-sectional area in the nozzle axial direction is the location where the nozzle's cross-sectional area is 1.1 times or less of the smallest cross-sectional area in the nozzle axial direction.

[0042] It should be noted that, in the top-blowing spray gun of the present invention configured as described above, a more preferred solution is as follows:

[0043] (1) The axial length of the above-mentioned spray nozzle at the above-mentioned spray outlet is less than 0.25 times the nozzle inner diameter corresponding to the minimum cross-sectional area of ​​the above-mentioned spray nozzle.

[0044] (2) As a jet nozzle, a straight pipe nozzle with a straight pipe section having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, or a Laval nozzle with an end enlargement in the throat having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, is used.

[0045] According to the present invention, by controlling the control gas ejected from the nozzle side near the portion of the nozzle with the smallest cross-sectional area in the longitudinal direction in the oxygen-containing gas injection nozzle of the top-blown lance without using a mechanically movable part, the gas flow rate can be controlled circumferentially without relying on the total gas flow rate. Therefore, even under the operating conditions of oxygen-feeding refining where molten iron or other materials splatter violently, operation can be performed without causing malfunction of the mechanically movable part. Furthermore, since the gas flow rate can be effectively increased even under insufficient expansion conditions at low gas flow rates, a top-blown oxygen-feeding method and the top-blown lance used therein can be realized with a wide range of variable gas flow rates. That is, even with a nozzle with a larger minimum inner diameter suitable for reducing spitting under high gas flow conditions, oxygen-feeding refining can be performed while suppressing the decrease in gas flow rate under low gas flow conditions. Attached Figure Description

[0046] Figure 1 This is a schematic diagram of a longitudinal section showing an example of a gas jet nozzle used in the top-blowing spray gun of the present invention.

[0047] Figure 2 In the text, (a) to (d) represent the characters used for... Figure 1 The diagram shows a cross-section of the throat at the gas outlet of the gas injection nozzle, used to illustrate the control of the gas injection nozzle.

[0048] Figure 3 To indicate Figure 2 The diagrams (a) to (d) show the behavior of the increase in jet velocity in the gas jet nozzle caused by the control gas flow rate.

[0049] Figure 4 This is a graph showing the result of adjusting the jet velocity at the maximum control gas flow rate ratio in the gas jet nozzle used in the top-blowing spray gun of the present invention, with the diameter of the control gas outlet × the number of control gas outlets / the throat diameter of the nozzle as the horizontal axis.

[0050] Figure 5 This is a graph showing the result of adjusting the jet velocity at the maximum control gas flow ratio in the gas jet nozzle used in the top-blowing spray gun of the present invention, with the gap of the slits / throat diameter of the jet nozzle as the horizontal axis.

[0051] Figure 6 This is a graph showing the relationship between the concentration of decarbonized carbon at the end of the decarburization process in the gas injection nozzle used in the top-blown spray gun of the present invention and the concentration of T.Fe in the slag (mass%).

[0052] Figure 7 This is a graph showing whether slag spraying occurs due to the control gas flow rate ratio during the initial stage of decarburization blowing using the present invention.

[0053] Figure 8 This is a graph showing the relationship between the control gas flow rate ratio and the furnace dust generation rate under the condition that the silicon concentration of molten iron is less than 0.4% by mass during decarburization blowing using the present invention.

[0054] Figure 9 This is a graph showing the relationship between the T.Fe concentration (mass%) in the slag and the control gas flow rate ratio when decarburization is carried out using the decarburization blowing process of the present invention until the carbon concentration is about 0.05% by mass.

[0055] Symbol Explanation

[0056] 1. Throat

[0057] 2. End enlargement section

[0058] 3. Spray outlet

[0059] 4. Gas storage tank Detailed Implementation

[0060] Hereinafter, embodiments of the present invention will be described with reference to the accompanying drawings.

[0061] Figure 1 This is a schematic diagram of the longitudinal section of a nozzle, representing an example of a gas injection nozzle for a top-blown spray gun used in this invention. Oxygen-containing gas for oxygen refining is injected from the gas storage tank 4 of the top-blown spray gun through an injection nozzle that penetrates the outer casing of the top-blown spray gun onto the bath surface. Figure 1 and Figure 2 In the examples shown in (a) to (d), for the sake of simplicity, only the front end of the top-blowing spray gun with a single nozzle is shown; the cooling water flow path of the water-cooled top-blowing spray gun housing is omitted. Here, industrial pure oxygen is generally used as the oxygen-containing gas, and sometimes a mixture of pure oxygen and nitrogen or argon is used, depending on the purpose.

[0062] Figure 1The Laval nozzle shown has a throat 1 with the smallest cross-sectional area in the axial direction of the nozzle and an extended end 2 adjacent to its downstream side. Sometimes, it also has a pointed tip (not shown) extending upstream of the throat 1, forming a pointed end-enlarged nozzle shape that guides the main supply gas into the throat 1. The top-blowing spray gun used in this invention has a gas injection nozzle with a control gas outlet 3, which is located on the nozzle side near the portion of the nozzle with the smallest cross-sectional area in the axial direction of the nozzle, and is configured such that at least a portion of the outlet exists in both spaces when bisected by any plane passing through the central axis of the nozzle. Control gas, whose flow rate can be independently controlled from the main supply gas supplied from the nozzle inlet, can be ejected into the nozzle from the control gas outlet 3, while oxygen-containing gas, which is the main supply gas, is supplied from the nozzle inlet side.

[0063] Here, the cross-sectional area of ​​the spray nozzle including the nozzle outlet 3 refers to the area enclosed by the nozzle outlet 3 in a plane perpendicular to the central axis of the spray nozzle, where the portion of the nozzle outlet 3 that does not actually have a spray nozzle side is interpolated with a smooth curved surface that is continuous with the nozzle side around the nozzle outlet 3. The resulting curved surface is taken as the side of the imaginary nozzle.

[0064] At this time, when the side surface of the jet nozzle, excluding the portion of the multiple nozzle outlets 3, is formed as the side surface of a rotating body centered on the central axis of the jet nozzle, the imaginary nozzle surface is equal to the side surface of this rotating body. When it is a Laval nozzle, the surface interpolated to the portion of the nozzle outlet 3 is mostly composed of a part of the side surface of a cylinder or cone or a combination thereof. However, if it also includes cases where the shape of the end enlargement 2 is not a truncated cone bell shape, or cases where the cross-sectional shape of the jet nozzle is not circular, then it is not necessarily limited to a part of the side surface of a cylinder or cone or a combination thereof.

[0065] Furthermore, as described later, when the nozzle 3 is formed in a slit shape around the entire circumference of the injection nozzle, the imaginary nozzle surface is obtained by interpolating the portion of the nozzle 3 with a smooth curve (including straight lines) that is continuous with the side of the nearby nozzle in a cross section containing the central axis of the injection nozzle.

[0066] It is known that in a top-blown spray gun without nozzle 3 and with a typical top-blown oxygen nozzle, the relationship between the oxygen flow rate and the pressure at the throat inlet is empirically expressed as follows (2):

[0067] Pt = Fo2 / (0.456 × n × dt) 2 (2)

[0068] Here, Pt is the gas pressure (absolute pressure) at the inlet of throat 1 (kgf / cm). 2 Fo2 is the oxygen flow rate (Nm³) injected from the top-blown nozzle. 3 / hr), n is the number of nozzles in the top-blowing spray gun, and dt is the inner diameter of the throat of the nozzle.

[0069] According to equation (2), the gas pressure Pt at the inlet of throat 1 is directly proportional to the gas flow rate and inversely proportional to the cross-sectional area of ​​throat 1 (or, Pt is directly proportional to the linear velocity of the gas (Nm / s)). The gas jet ejected from the nozzle is fundamentally powered by this gas pressure Pt, and qualitatively there is a trend that the higher the gas pressure Pt, the higher the velocity or kinetic energy of the gas jet.

[0070] In contrast, if the total gas flow rate ejected from the nozzle is constant before the control gas is ejected from the outlet 3, a region with low axial mass flow velocity is generated near the outlet 3 of the throat 1. In other regions of the throat 1's cross-section (the section perpendicular to the central axis of the nozzle), the mass flow velocity (mass flow rate per unit area) increases compared to the case where no control gas is ejected. Therefore, it is observed that the gas pressure of the main supply gas increases at the inlet of the throat 1, and the velocity of the gas jet ejected from the nozzle increases. This phenomenon can also be described as an effect of seemingly reducing the cross-sectional area of ​​the throat 1, but it is significant even if the proportion of the control gas relative to the main supply gas is small. This phenomenon is observed not only in the case of a Laval nozzle with a control gas outlet 3 in the throat 1, but also in a straight pipe nozzle with a fixed axial cross-sectional area where a control gas outlet is provided at a certain axial position. In a straight pipe nozzle without an end enlargement 2, the nozzle can be positioned at any axial position as long as the axial position of the nozzle with multiple outlets 3 is the same relative to all outlets 3. That is, the nozzle outlet 3 is located on the side of the nozzle where the cross-sectional area of ​​the nozzle is the smallest in the axial direction of the nozzle.

[0071] In order to efficiently convert the increase in gas pressure of the main supply gas at the throat inlet caused by the introduction of control gas from the nozzle 3 into kinetic energy and increase the jet velocity, it is necessary to consider the influence of nozzle shape in the same way as in the case of a typical Laval nozzle. The inventors have found that under specific nozzle shape conditions, a particularly good effect of increasing jet velocity can be obtained. That is, under the condition that the gas pressure at the throat inlet of the main supply gas is higher than the appropriate expansion pressure Po determined by the following formula (1) for the opening ratio (Ae / At) of the injection nozzle, and the apparent underexpansion condition, the jet velocity can be increased more effectively than when this condition is not met:

[0072] Ae / At = (5 5／2 / 63 )×(Pe / Po) －5／7 ×[1-(Pe / Po] 2／7 ] －1／2 ···(1)

[0073] Here, At: minimum cross-sectional area of ​​the injection nozzle (mm²) 2 Ae: Exit cross-sectional area of ​​the injection nozzle (mm²) 2 Pe: Atmospheric pressure at the nozzle outlet (kPa), Po: Appropriate expansion pressure of the nozzle (kPa). The effect of nozzle shape on increasing the jet velocity can be explained as follows.

[0074] That is, in a typical Laval nozzle, when the gas pressure at the inlet of the throat 1 is higher than the appropriate expansion pressure, the expansion portion 2 at the end of the Laval nozzle becomes under-expanded. The gas is ejected from the nozzle outlet under high pressure and expands outside the nozzle with the shock wave, thus generating energy loss. Compared with the case where the nozzle expands appropriately under the same gas pressure at the inlet of the throat 1 and has a larger opening, the jet velocity is reduced.

[0075] In contrast, when control gas is ejected from the multiple nozzle outlets 3 on the side of the nozzle of the throat 1 (or the straight pipe section where the cross-sectional area of ​​the nozzle is smallest in the nozzle axial direction), the gas boundary layer of the main supply gas formed along the side (wall) of the nozzle of the throat 1 is stripped away from the side of the nozzle, resulting in an apparent reduction in the nozzle cross-sectional area of ​​the throat 1. On the other hand, it is believed that the effect of reducing the nozzle cross-sectional area is relatively reduced by accelerating the control gas in the gas injection direction of the injection nozzle at the nozzle outlet. Therefore, by introducing control gas, an effect of substantially increasing the opening ratio compared to the actual nozzle shape is achieved, resulting in a substantial expansion at the gas pressure at the inlet of the throat 1 that is higher than the appropriate expansion pressure determined by the above formula (1) according to the nozzle shape (opening ratio), and the jet velocity increases. In addition, when using a nozzle with an opening ratio determined by the above formula (1) for the gas pressure at the inlet of the throat 1, it becomes substantially over-expanded, resulting in energy loss. In this way, when the control gas is ejected from the multiple nozzle outlets on the side of the nozzle of the throat 1 (or the part of the nozzle with the smallest cross-sectional area in the nozzle axis), the gas pressure of the main supply gas at the inlet of the throat 1 is higher than the appropriate expansion pressure Po determined by the following formula (1) according to the shape (opening ratio) of the injection nozzle, and the apparent expansion is insufficient, the jet flow velocity can be effectively increased compared with the case where this condition is not met.

[0076] To confirm the function of increasing the jet velocity obtained from the control gas as described above, the following was used: Figure 1Model experiments were conducted using nozzles of approximate shapes as shown, to investigate the effect of the control gas on the jet velocity. The shape conditions of the nozzles used are shown in Table 1. Nozzles A1-A3 and B are Laval nozzles with a throat 1, and nozzles C1-C6 are straight pipe nozzles with control gas outlets at a specified distance from the nozzle outlet. The control gas outlet is as follows under all conditions... Figure 2 As shown in the cross-sectional view of the throat of the injection nozzle in (c), eight inlet holes (control gas inlet holes) are evenly arranged circumferentially, forming open ends with an inner diameter of 1 mm. C5 and C6 close four of the eight nozzle outlets, with C5 making the four outlets adjacent to each other and C6 making the outlets spaced one apart. The area ratio of the control gas outlets in Table 1 refers to the ratio of the total cross-sectional area of ​​the control gas inlet holes to the minimum cross-sectional area of ​​each nozzle.

[0077] [Table 1]

[0078]

[0079] The main supply gas and high-pressure air used as control gas were supplied at the flow rates shown in Table 2. The results of measuring the jet velocity on the central axis 200 mm from the nozzle tip, as well as the supply pressures of the main supply gas and control gas, are shown in Table 2. In this experiment, the total gas flow rate (the sum of the control gas flow rate and the main supply gas flow rate) was varied for each nozzle within three conditions. The study compared the case where no control gas was supplied with the case where the ratio of the control gas flow rate to the total gas flow rate was 20%. It should be noted that the main shapes of the nozzles used in the model test, such as the minimum diameter and opening ratio, shown in Table 1, were determined to be similar to the gas injection nozzle of the top-blowing spray gun used in the actual 300t scale described later. In addition, the gas flow rates in the model test shown in Table 2 were set to approximately 1 / 100 of the operating condition range of the gas injection nozzle of the actual machine, so that the gas pressure or linear velocity was similar to the operating conditions of the actual machine.

[0080] [Table 2]

[0081]

[0082] The jet gas velocity difference in Table 2 is the difference in jet gas velocity caused by the presence or absence of control gas, assuming the same nozzle shape and total gas flow rate. According to the results in Table 2, even with a constant total gas flow rate, the pressure of the main supply gas can be increased by ejecting control gas, thereby increasing the jet velocity. It is evident that the increase in jet velocity is particularly significant when the pressure of the main supply gas exceeds the appropriate expansion pressure of each nozzle. This is attributed to the apparent increase in the opening ratio created by ejecting control gas, as described above, which brings the conditions closer to appropriate expansion.

[0083] Furthermore, it is known that regardless of the type of Laval nozzle or straight nozzle, an amplification effect can be achieved as long as an outlet is present on the side of the nozzle at the location where the nozzle cross-sectional area is minimized (examples A1, B, and C1-C6) or in a nearby location (examples A2 and A3). Moreover, it is considered that since no effect is obtained when the control gas is ejected from the nozzle in one direction, the control gas outlet should be configured such that at least a portion of the outlet exists in both spaces when the nozzle is bisected by any plane passing through the central axis of the nozzle.

[0084] Here, referring to A1 to A3 of the Laval nozzles using Tables 1 and 2, we investigate the "location with the smallest nozzle cross-sectional area". First, it is found that the location with the nozzle outlet in A1, due to the enlargement length of 4 mm and the distance between the control gas outlet and the nozzle outlet of 4 mm, is the throat 1 with the smallest nozzle cross-sectional area in the nozzle axial direction. Furthermore, it is found that the location with the nozzle outlet in A2, due to the enlargement length of 4 mm and the distance between the control gas outlet and the nozzle outlet of 2.7 mm, is the location with a nozzle cross-sectional area 1.06 times the smallest in the nozzle axial direction. Additionally, the location with the nozzle outlet in A3, due to the enlargement length of 4 mm and the distance between the control gas outlet and the nozzle outlet of 2 mm, is the location with a nozzle cross-sectional area 1.14 times the smallest in the nozzle axial direction. Under these conditions, the "jet gas velocity difference (m / s)" in Table 2 for A1 to A3 is calculated with control gas and a total gas flow rate of 1.1 Nm³. 3 When comparing the values ​​per minute, nozzle A1, with a ratio of "1" to the minimum cross-sectional area, has a speed of +20 m / s; nozzle A2, with a ratio of "1.06" to the minimum cross-sectional area, has a speed of +10 m / s; and nozzle A3, with a ratio of "1.14" to the minimum cross-sectional area, has a speed of +0. Therefore, it can be understood that in this invention, when using a Laval nozzle, the portion near the location with the minimum cross-sectional area refers to a portion where the nozzle's cross-sectional area is preferably 1.1 times or less of the minimum cross-sectional area along the nozzle's axial direction.

[0085] Next, the supply conditions for the control gas will be explained.

[0086] In a jet nozzle with the same Laval nozzle shape as nozzle B in Table 1, the effect of the control gas flow rate ratio (the ratio of control gas flow rate to total gas flow rate) on the jet velocity was investigated under various modifications to the control gas outlet. Here, as... Figure 2 As shown in (a) to (d), the control gas outlets are arranged in a manner where 2, 4, or 8 outlets are evenly distributed circumferentially, or distributed throughout the entire circumference, and are formed in a slit shape, thus being rotationally symmetrical with respect to the central axis of the injection nozzle. When multiple outlets are provided, the outlet of each injection nozzle is formed as the open end of a control gas inlet hole with a circular cross-section and an inner diameter of 1 mm. Furthermore, when the outlets are slit-shaped, the width of the slit gap is 1 mm. The total gas flow rate in each injection nozzle is 1.1 Nm³. 3 The flow rate of the control gas was varied within the range of 0-30% at a constant rate of / min, and the jet velocity at the center axis 200mm from the nozzle tip was measured. The results of the jet velocity measurement are shown below. Figure 3 .like Figure 3 As shown, even if the control gas nozzle is a slit covering the entire circumference, and even when multiple nozzles are configured, the jet flow rate effect is still achieved. It can be said that in order to achieve, to some extent, the effect of apparently reducing the nozzle cross-sectional area at the throat, the control gas flow rate ratio is preferably 5% or more. Furthermore, there is no particular upper limit to the control gas flow rate ratio, but to avoid enlarging the control gas flow path and control gas supply system, it is preferably 50% or less, more preferably 30% or less.

[0087] Furthermore, it can be known that in Figure 3 Among all the nozzles shown, there exists a control gas flow rate ratio that maximizes the jet velocity. However, if the control gas flow rate ratio is increased beyond this ratio, a gradual decrease in jet velocity is sometimes observed. This is believed to be due to the relationship between the effect of substantially increasing the opening ratio compared to the actual nozzle shape caused by the introduction of control gas and the effect of increasing the pressure of the main supply gas at the throat inlet, resulting in a control gas flow rate ratio that is essentially appropriately expanded.

[0088] Next, in a spray nozzle with the same Laval nozzle shape as nozzle B in Table 1, the control gas outlet was formed as the open end of 2 to 8 control gas inlet holes with circular cross-sections evenly arranged in the circumferential direction. The inner diameter of the control gas inlet holes was varied from 0.8 to 1.2 mm, and the jet velocity was measured to investigate how the proportion of the area containing the control gas outlet affects the circumferential direction of the throat. In each nozzle, the total gas flow rate was 1.1 Nm³. 3 Under certain conditions, the jet velocity at the control gas flow rate ratio with the control gas outlet diameter × number of control gas outlets / nozzle throat diameter as the horizontal axis is adjusted, and the results are shown in the figure. Figure 4 .

[0089] according to Figure 4 From the viewpoint of the apparent reduction in the nozzle cross-sectional area at the throat, it is preferable that the proportion of the area where the nozzle outlet is located circumferentially in the throat (or the straight section where the nozzle cross-sectional area is smallest in the nozzle axial direction) is relatively large. Therefore, the circumferential outlets provided on the side of the injection nozzle in multiple directions are preferably such that the product of the total circumferential extension of the side of the injection nozzle (the diameter in the direction perpendicular to the central axis of the injection nozzle and the central axis of the control gas inlet orifice, or the diameter of the inlet orifice into which the control gas is introduced to the outlet) and the diameter of the outlet, i.e., the diameter of the outlet and the number n of outlets per injection nozzle, is at least 0.4 times the throat diameter of the injection nozzle or the inner diameter of the nozzle at the part where the cross-sectional area is smallest.

[0090] Furthermore, in a Laval nozzle with the same shape as nozzle B in Table 1, the control gas outlet is a slit covering the entire circumference of the nozzle. The jet velocity is measured in the same manner as described above, with the slit gap interval varying from 0.6 mm to 2.0 mm. For each nozzle, the jet velocity at the control gas flow rate ratio with the slit gap interval / nozzle throat diameter as the horizontal axis is calculated, and the results are shown below. Figure 5 .

[0091] according to Figure 5It is known that if the nozzle outlet is provided in a slit shape along the entire circumference of the side of the nozzle, the axial length of the nozzle with the slit-shaped gap becomes too large, and the effect of increasing the jet velocity tends to decrease. Therefore, the axial length of the nozzle with the slit-shaped nozzle outlet is preferably less than 0.25 times the inner diameter of the nozzle at the part with the smallest cross-sectional area. Furthermore, if the slit-shaped gap is excessively larger than 0.25 times the inner diameter of the nozzle, the flow rate of the control gas required to achieve the aforementioned effect of seemingly reducing the nozzle cross-sectional area at the throat increases, necessitating a larger control gas flow path and control gas supply system, which is also not preferable from this perspective.

[0092] In addition, when describing the characteristics of the nozzle, such as Figure 2 As shown in the cross-sectional views of the throat in (a) to (d), there need to be two or more nozzles, or they can be slits covering the entire circumference of the nozzle. If the nozzles are arranged asymmetrically with respect to the central axis of the nozzle, as described in Patent Document 3, there is a tendency for the gas jet ejected from the nozzle to deflect from the central axis. Therefore, it is preferable that the nozzles are arranged such that at least a portion of the nozzle is present in both spaces when the nozzle is bisected by any plane passing through the central axis of the nozzle. In this case, from the viewpoint of the effect of seemingly reducing the cross-sectional area of ​​the nozzle at the throat, it is preferable that the multiple nozzles are all in the same position along the axial direction of the nozzle, and it is not necessary to strictly align the positions along the axial direction of the nozzle. If the nozzles are arranged close to each other along the axial direction of the nozzle, and at least a portion of the nozzle is present in both spaces when the nozzle is bisected by any plane passing through the central axis of the nozzle, although the efficiency is lower than that of all nozzles being arranged in the same position along the axial direction of the nozzle, a similar effect of increasing the jet velocity is obtained.

[0093] When the outlets of the control gas are arranged circumferentially along the sides of the nozzle in multiple directions, the guide paths for the control gas to be introduced into the multiple outlets are interconnected within the top-blown spray gun. This simplifies the flow control system and supply path of the control gas and ensures a balanced supply of control gas from each outlet. More preferably, the guide paths for introducing the control gas into the multiple outlets are arranged in an annular gas flow path around the spray nozzle.

[0094] In addition, it is preferable that the entire nozzle outlet is included in the throat. Sometimes the length of the throat is short and smaller than the diameter of the nozzle axial direction of the nozzle outlet. Even if a part of the nozzle outlet is included in the downstream end enlargement or the upstream tip detail (not shown), as long as the center position of the nozzle outlet is included in the throat, or the entire throat is included in the range of the nozzle axial direction of the nozzle outlet, the function of controlling the jet flow velocity described later will not be significantly different, and the same effect will be obtained.

[0095] Furthermore, the effect of seemingly reducing the nozzle cross-sectional area by ejecting control gas from the side of the nozzle is not necessarily limited to the case where the cross-sectional area of ​​the nozzle outlet is strictly at its minimum along the nozzle axial direction. It simply means that the effect of increasing the jet velocity is most effective at that location. Even at locations close to the minimum cross-sectional area along the nozzle axial direction, a similar increase in jet velocity can sometimes be achieved. However, if the cross-sectional area of ​​the nozzle at the axial position of the nozzle outlet becomes larger, a larger amount of control gas is required, and the efficiency of increasing jet velocity decreases. Therefore, it is preferable to place the nozzle at a cross-sectional area of ​​1.1 times or less of the minimum cross-sectional area.

[0096] Furthermore, in order to more effectively achieve the aforementioned effect of seemingly reducing the nozzle cross-sectional area at the throat, it is preferable that the linear velocity (Nm / s) of the control gas ejected into the injection nozzle is relatively large. If it is within approximately 1 / 2 to 2 times the linear velocity of the main supply gas at the throat (the average value of the entire cross-section of the throat), the pressure of the control gas will not become excessively high, thus effectively achieving the effect of seemingly reducing the nozzle cross-sectional area at the throat, which is therefore preferable. Based on the model test results shown above, regarding the preferred conditions for increasing the jet velocity through the control gas, even when the scale or size of dimensionless indicators such as flow ratio, length ratio, area ratio, and linear velocity ratio differ significantly, including in actual machine cases, as long as the range of gas pressure or linear velocity at the nozzle is the same, it is sufficiently effective, and the preferred range of the corresponding dimensionless indicators can be directly adopted.

[0097] Next, the inventors conducted in-depth research on a method for controlling the flow rate or dynamic pressure of the jet by using the top-blown lance of the present invention to achieve stable operation in oxygen refining processes such as decarburization blowing in a converter, while reducing the amount of furnace dust generated and the oxidation loss of iron.

[0098] Generally, oxygen refining of iron and steel is carried out for the purposes of desiliconization, decarburization, and dephosphorization. However, in the initial stage of refining, the goal is to increase the oxygen supply rate to effectively remove impurity elements. In the later stage of refining, the concentration of impurity elements decreases, and reactions other than those intended for this purpose, such as the formation of iron oxide, become dominant. Therefore, oxygen supply methods that reduce the oxygen supply rate are mostly chosen. When oxygen is supplied from a top-blown lance, the kinetic energy of the top-blown oxygen jet changes with this change in oxygen supply rate. Therefore, the collision state of the top-blown oxygen jet with the molten slag and the hot metal bath surface may change, thus affecting the reaction rate.

[0099] For example, in the decarburization refining of molten iron, if the supply rate of top-blown oxygen is reduced at the end of the oxygen refining process to suppress the formation of iron oxide, the kinetic energy of the top-blown oxygen jet decreases. Changes in the stirring and mixing state at the collision point (ignition point) of the top-blown oxygen jet lead to a decrease in decarburization efficiency. Therefore, in such cases, reducing the lance height is used to suppress the decrease in the kinetic energy of the top-blown oxygen jet. However, there is a limit to the safe and feasible lance height, making it difficult to fully address this issue.

[0100] In the oxygen-feeding refining method for molten iron of the present invention, under such circumstances, by adjusting the supply rate of the control gas according to the supply rate of oxygen-containing gas blown from the top-blown lance to the molten iron, the kinetic energy of the top-blown oxygen jet can be increased, thus increasing the degree of freedom in obtaining refining conditions with an effective reaction rate. For example, in the decarburization refining of molten iron, when the top-blown oxygen supply rate is reduced at the end of oxygen-feeding refining after 85% of the total oxygen supply, by simultaneously spraying the control gas and supplying oxygen as the main supply gas, the decrease in decarburization efficiency can be suppressed, and the formation of iron oxide can be suppressed more effectively. At this time, by not supplying the control gas in the refining stages other than the final stage, excessive molten iron scattering and furnace dust generation can be suppressed even in the earlier refining stages when the oxygen supply rate is high. By changing the supply rate of the control gas along with the oxygen-feeding refining, effective refining conditions can be maintained overall.

[0101] To verify the effect of increasing the jet velocity at the molten iron bath surface and suppressing iron oxide formation under the same total gas flow rate and lance height conditions by supplying control gas, a 2-ton top-and-bottom blown refining furnace was used for molten iron decarburization treatment, and the influence of control gas on the iron oxide concentration in the slag was investigated. It was concluded that in refining tests using a small furnace, by ensuring that the oxygen supply per unit mass of molten iron, the amount and rate of refining agent supply, and the stirring power density (W / t) based on the bottom-blown gas are identical to those in the actual process, tests simulating the refining reaction in the actual process can be conducted. Under the oxygen flow rate determined thereby, the top-blown lance was designed to achieve a range of gas pressure or linear velocity at the nozzle that is identical to that in the model tests of the top-blown lance or the aforementioned injection nozzle in the actual process. Furthermore, regarding the lance height, an empirical formula for calculating the molten iron depression depth was used, and the ratio of depression depth to iron bath depth was determined to be identical to the operating range of the actual process.

[0102] As shown in Table 3, the top-blown lances used in the experiment employed two types of top-blown lances: a single-hole lance D with a straight-tube-type nozzle and a five-hole lance E. Each lance had four control gas outlets arranged in each nozzle in a rotationally symmetrical manner with respect to the central axis of each nozzle. According to the main experimental conditions shown in Table 4, decarburization was carried out under a constant total oxygen flow rate while bottom-blowing a small amount of argon gas to agitate the molten iron until a low carbon concentration region was reached. For each top-blown lance, the condition of no control gas supply and the condition of supplying approximately 23% of the total oxygen flow rate as control gas were compared. The relationship between the decarburization carbon concentration (mass%) at the end of the decarburization process and the T.Fe concentration (mass%) in the slag was measured, and the results are shown in Table 5. Figure 6 .

[0103] [Table 3]

[0104]

[0105] [Table 4]

[0106]

[0107] [Table 5]

[0108]

[0109] According to Table 5 and Figure 6 The results shown demonstrate that by injecting control gas from the control gas outlet, compared to existing techniques without control gas, even under the same total gas flow rate and lance height, the amount of iron oxide (T.Fe) in the slag is relatively reduced, thus suppressing iron oxidation loss. This is attributed to the increased flow velocity of the oxygen jet colliding with the iron bath, resulting in enhanced stirring force at the ignition point, due to the effect of the control gas. In this experiment, control gas was supplied throughout the entire blowing process. However, it is known that the increase in iron oxide concentration in the slag during decarburization refining is significant towards the end of refining. For example, it was demonstrated that even when control gas is supplied only at the end of oxygen refining (after 85% of the total oxygen supply), the same effect of suppressing iron oxidation loss is achieved, and varying the supply rate of control gas as oxygen refining progresses is effective.

[0110] In addition, it is also effective to change the supply rate of control gas based on the results of detecting the refining state in oxygen refining. For example, it is effective to change the supply rate of control gas to adjust the iron oxide formation rate based on the detection of slag foaming height or the results of measuring the decarbonization oxygen efficiency over time according to the analysis information of the exhaust gas (for example, in the case of excessive iron oxide concentration in the slag, in order to reduce the iron oxide formation rate, the supply of control gas is started to increase the dynamic pressure of the top blown oxygen jet).

[0111] Furthermore, adjusting the supply rate of the control gas based on refining conditions such as the temperature of the molten iron, silicon concentration, carbon concentration, and amount of iron filings used, as determined before the start of oxygen refining, is also effective. For example, in the decarburization refining of molten iron with a silicon concentration of 0.40% by mass or higher before the start of oxygen refining, in the initial stage of oxygen refining when the total oxygen content of the oxygen-containing gas to be supplied is less than 20%, there is a tendency for slag spraying to occur under refining conditions of high oxygen supply rate and high lance height. In this case, the following method can be adopted: by simultaneously spraying the control gas and supplying the oxygen-containing gas as the main supply gas, the dynamic pressure of the top-blown oxygen jet is increased to suppress the formation of excessive iron oxide and prevent slag spraying. At the same time, in the decarburization refining of molten iron with a silicon concentration of less than 0.40% by mass before the start of oxygen refining, the control gas is not supplied in the initial stage of oxygen refining, so that the dynamic pressure of the top-blown oxygen jet is pushed to a lower level, thereby suppressing the scattering of molten iron and the generation of furnace dust.

[0112] In the decarburization blowing process of a converter, it is known that when the silicon concentration in the molten iron is high before blowing, slag spraying, sometimes known as slag ejection, sometimes occurs. This is because if a large amount of silica is generated in the early stages of blowing, during which almost no CaO-based solvents such as quicklime are dissolved into the liquid slag (slag formation), the apparent volume of CO bubbles generated during the decarburization reaction is rapidly increased by about 10 times in the large amount of highly viscous molten slag (slag bubbling). In particular, as the slag bubbling increases in thickness, there is a decrease in the intensity of the top-blown oxygen jet and a change in the collision between the oxygen jet and the molten iron and slag. The proportion of oxygen consumed by the oxidation of iron increases, leading to a tendency for the iron oxide concentration in the slag to rise. If the iron oxide concentration in the slag increases, the reaction with carbon in the molten iron droplets in the molten iron bath and slag causes the tiny CO bubbles formed in the slag to increase in size, further promoting bubbling. Therefore, bubbling is accelerated and sometimes leads to slag spraying.

[0113] As a method to prevent slag spraying, it was considered to reduce the lance height based on the foaming height of the slag and to ensure the dynamic pressure of the top-blown jet colliding with the molten iron bath to suppress the formation of excessive iron oxide. However, reducing the lance height under the high oxygen supply rate conditions in the early stages of blowing would lead to lance damage due to flying molten iron, increased repair frequency, or high risk of operational difficulties due to water leakage, which is not a good solution. Slag spraying has become a major factor seriously hindering operation, so slag spraying is usually suppressed by keeping the oxygen supply rate at a lower level in the early stages of blowing when the silicon concentration in the molten iron before blowing is high. However, reducing the oxygen supply rate leads to a prolongation of blowing time. Therefore, the inventors investigated the influence of the silicon concentration in the molten iron before blowing and the ratio of the control gas flow rate supplied to the nozzle on slag spraying without reducing the oxygen supply rate in the early stages of blowing.

[0114] In a 2-ton top-and-bottom blown refining furnace, decarburization was performed on molten iron with various silicon concentrations. The effects of the control gas on slag generation, furnace dust generation, and T.Fe concentration in the slag were investigated. The basic experimental conditions, except for the control gas flow rate, were the same as those shown in Table 4. The silicon concentration of the molten iron before decarburization varied from 0.1% to 0.5% by mass. The same top-blown lance as lance E in Table 3 was used. Under the condition of a constant total oxygen flow rate, various changes were made to the control gas flow rate ratio to achieve a low carbon concentration of approximately 0.05% by mass.

[0115] The results regarding whether slag spraying occurs during the decarburization blowing of molten iron with a silicon concentration of 0.4% or higher before blowing are due to the control gas flow rate ratio in the initial blowing stage are presented below. Figure 7 It should be noted that no slag spraying was observed during the decarburization blowing of molten iron with a silicon concentration of less than 0.4% by mass before blowing. Based on these results, it can be concluded that during the decarburization blowing of molten iron with a silicon concentration of 0.4% or more before blowing, slag spraying during the initial blowing stage can be suppressed by supplying control gas under appropriate conditions from the control gas outlet of the oxygen injection nozzle located in the top blowing lance.

[0116] Furthermore, the relationship between the control gas flow rate ratio and the furnace dust generation rate under the condition that the silicon concentration of molten iron is less than 0.4% by mass is shown in the figure. Figure 8 It is known that increasing the flow rate ratio of the control gas tends to increase the rate of furnace dust generation. It is known that the main component of furnace dust in decarburization refining is generated by tiny droplets (bubble rupture) produced as CO bubbles break, with the generation rate being particularly high from the initial stage to the peak of decarburization. It is believed that if the flow rate of the oxygen jet increases due to the supply of control gas, the physically dispersed molten iron droplets increase, thereby increasing the rate of furnace dust generation due to secondary bubble rupture, or increasing the proportion of furnace dust carried out of the furnace by the increased gas flow rate, thus increasing the rate of furnace dust generation. Furthermore, in the decarburization of molten iron with a low silicon concentration after pretreatment, the amount of cover slag generated is small, thus the rate of furnace dust generation tends to increase. Therefore, in the decarburization of molten iron with a silicon concentration of less than 0.4% by mass, it is preferable to avoid an increase in the rate of furnace dust generation by performing blowing without supplying control gas during the peak of decarburization.

[0117] The relationship between the T.Fe concentration (mass%) in the slag during decarburization blowing in molten iron with a silicon concentration of less than 0.4% by mass until the carbon concentration is approximately 0.05% by mass and the ratio of the control gas flow rate is shown in the figure. Figure 9It can be seen that by supplying control gas under appropriate conditions, the T.Fe in the slag can be reduced, thereby suppressing the oxidation loss of iron. This is believed to be because the same trend exists in the decarburization treatment of molten iron with a silicon concentration of 0.4% by mass or higher, where the effect of the control gas increases the flow rate of the oxygen jet and enhances the stirring force at the ignition point.

[0118] Based on the above insights, the following refining method can be preferred: In the decarburization process of molten iron with a silicon concentration of 0.4% by mass or more, during the initial stage of oxygen refining (before 20% of the total oxygen is supplied) and the final stage of oxygen refining (after 85% of the total oxygen is supplied), control gas is supplied under appropriate conditions from the control gas outlet of the oxygen injection nozzle installed in the top blow lance, thereby relatively increasing the flow rate of the oxygen jet. Control gas is not supplied during other periods.

[0119] In addition, the preferred refining method is as follows: during the decarburization process of molten iron with a silicon concentration of less than 0.4% by mass, at the end of the oxygen refining process after 85% of the total oxygen supply, control gas is supplied under appropriate conditions from the control gas outlet of the oxygen injection nozzle installed in the top blow lance, thereby increasing the flow rate of the oxygen jet relatively, and control gas is not supplied during other periods.

[0120] Example

[0121] The following describes a practical example of applying the oxygen-feeding refining method for molten iron of the present invention to industrial-scale converter decarburization treatment.

[0122] In a 300t top-and-bottom blown converter, the design of the nozzles for various top-blown lances was modified to decarburize molten iron, and the impact on furnace dust generation, iron yield, and slag formation was investigated. Molten iron was received in a blast furnace in a mixing car pre-loaded with iron filings and transported to the ironmaking plant. A specified amount of molten iron was then poured into a ladle, where a mechanically stirred desulfurization device was used for desulfurization. After the desulfurized slag was discharged from the ladle, molten iron was charged into a converter pre-loaded with approximately 30 tons of iron filings for further decarburization. The total amount of molten iron and iron filings charged in a single blowing operation was approximately 300 tons. The temperature of the molten iron when charged into the converter was 1280–1320℃, with a silicon concentration of 0.20–0.60% by mass and a carbon concentration ranging from 4.0% to 4.4% by mass.

[0123] Based on information such as the amount and temperature of the molten iron, silicon and carbon concentrations, the amount of iron filings, the target steel temperature, and carbon concentration, the total oxygen supply, heating materials, and cooling materials added during blowing are determined using static control. Additionally, the amount of by-products such as quicklime is determined to achieve a calculated basicity (CaO mass% / SiO2 mass%) of 3.5 in the decarburized slag, and the entire amount is added at the beginning of blowing. At this point, the slag production is adjusted as needed based on the phosphorus concentration of the target molten steel.

[0124] The total oxygen supply rate and lance height (the distance from the static bath surface of molten iron to the tip of the lance) during decarburization blowing are 750 Nm from the beginning to the middle of blowing, except for the end of the blowing process. 3 / min (2.5Nm) 3 / (min·t)) and 4.0m, respectively, at the end of the blowing process after supplying 85% of the total oxygen determined based on static control, 450Nm 3 / min (1.5Nm) 3 / (min·t)) and 2.5m. It should be noted that these lance heights were set based on past operational performance using lance F, with the corresponding total oxygen supply rate, as the lower limit of the lance height for stable operation where the damage condition of the top-blown lance does not vary significantly. Additionally, 30Nm is blown from the bottom through multiple blow plugs located at the bottom of the converter throughout the blowing process. 3 / min (0.10Nm) 3 Argon gas ( / (min·t)).

[0125] At the end of the blowing process, based on the temperature and carbon concentration of the molten steel measured using a sub-spray lance, the amount of oxygen supplied after the measurement and the amount of cooling material added are determined. Blowing ends at the point when the determined amount of oxygen is supplied, and the molten steel is poured into a ladle. Subsequently, the molten steel, after its composition and temperature have been adjusted by ladle refining using an RH degassing unit or a bubbling unit, is supplied to a continuous casting unit for continuous casting of slabs, etc.

[0126] The conditions for the eight top-blowing spray guns used in the experiment are shown in Table 6 below.

[0127] [Table 6]

[0128]

[0129] The spray gun F is a top-blown spray gun with a Laval nozzle used in previous operations. Spray guns G and H were modified from the nozzle shape of spray gun F to reduce iron scattering loss and furnace dust generation by lowering the jet flow velocity at high oxygen flow rates. In spray gun G, the throat diameter was increased to 66 mm, and in spray gun H, a straight-tube type nozzle with an inner diameter of 70 mm was used. It should be noted that, from the viewpoint of ensuring the water-cooling structure required for the top-blown spray gun, it is difficult to increase the outlet diameter of the nozzle to more than 70 mm.

[0130] Spray gun I is a top-blowing spray gun of this invention, located at the throat of each nozzle of spray gun G. Spray gun J is located at a position 70 mm from the outlet of each nozzle of spray gun H. Each nozzle has eight control gas inlet holes with open ends forming a circular cross-section of 10 mm in diameter, evenly arranged circumferentially on the inner surface of the nozzle. Spray guns K to M are also top-blowing spray guns of this invention, with different configurations of control gas outlets at a position 70 mm from the outlet of each nozzle of spray gun H. In spray guns K and M, slit-shaped control gas outlets with gaps of 3 mm and 10 mm, respectively, are provided around the entire circumference of the inner surface of each nozzle. In spray gun N, each nozzle has four control gas inlet holes with open ends forming a circular cross-section of 6 mm in diameter, evenly arranged circumferentially on the inner surface of the nozzle.

[0131] The inlet paths for the control gas from each nozzle of each spray gun to each control gas outlet are interconnected within the spray gun. Industrial pure oxygen, controlled to a specified flow rate, is supplied as the control gas from the control gas supply device. Regardless of the type of top-blown spray gun used, the control gas flow rate ratio (the ratio of control gas flow rate to total gas flow rate) shown in Table 6 is always used.

[0132] Next, the slag formation when using each top-blowing spray gun and the resulting operating methods will be explained.

[0133] In the case of lance F, no slag that would hinder operation is generated. In the case of lance G, if the silicon concentration of the molten iron is 0.50% by mass or more, and in the case of lance H, if the silicon concentration of the molten iron is 0.40% by mass or more, large slag is sometimes generated, making it difficult to maintain stable and continuous operation. Therefore, in operation using lance G, by pre-desiliconizing the molten iron in the mixing car and mixing it with molten iron with a low silicon concentration, the silicon concentration of the molten iron charged into the converter is limited to less than 0.50% by mass for continuous operation. In addition, in operation using lance I, which has the same nozzle shape as lance G, during the initial blowing stage (up to 20% of the total oxygen content determined based on static control), control gas is supplied when the silicon concentration of the molten iron charged into the converter is 0.50% by mass or more, and no control gas is supplied when the silicon concentration of the molten iron charged into the converter is less than 0.50% by mass. Furthermore, in operation using lance H, the silicon concentration of the molten iron charged into the converter is also limited to less than 0.40% by mass and operated continuously. Additionally, in operation using lances J to M, which have the same nozzle shape as lance H, during the initial blowing stage (up to 20% of the total oxygen determined based on static control), control gas is supplied when the silicon concentration of the molten iron charged into the converter is 0.40% by mass or higher; when the silicon concentration is less than 0.40% by mass, control gas is not supplied. At this time, the ratio of molten iron after pre-desiliconization treatment in operation using lance G, and the ratio of charge with a silicon concentration of 0.50% by mass or higher when charging into the converter in operation using lance I, are both approximately 10%.

[0134] Furthermore, in operation using lance H, the silicon concentration of the molten iron charged into the converter is also limited to less than 0.40% by mass and operated continuously. Additionally, in operation using lances J to M, which have the same nozzle shape as lance H, during the initial blowing stage (up to 20% of the total oxygen determined based on static control), control gas is supplied when the silicon concentration of the molten iron charged into the converter is 0.40% by mass or higher; when the silicon concentration is less than 0.40% by mass, control gas is not supplied. At this time, the ratio of molten iron after pre-desiliconization treatment in operation using lance H, and the ratio of charge with a silicon concentration of 0.40% by mass or higher when charging into the converter in operation using lances J to M, are both approximately 40%.

[0135] Furthermore, regardless of the type of spray gun with a control gas outlet used, the total oxygen supply rate is reduced and control gas is supplied simultaneously during the final stage of blowing, after 85% of the total oxygen amount determined based on static control is supplied. Additionally, during periods other than the initial and final stages of blowing, no control gas is supplied, regardless of the type of spray gun with a control gas outlet used.

[0136] Approximately 200 blow cycles were performed on each top-blown lance. The average results of the dust generation (in original units) and iron yield for each blow cycle are shown in Table 7 below. Dust generation is the average in original units calculated from the amount of dust generated during the use of each top-blown lance. Iron yield is calculated as the sum of the product quantity, sheet quantity, and raw material metal recovered for reuse throughout the process up to continuous casting. Additionally, Table 7 also shows the average back pressure (supply pressure of the main gas to the lance) of each lance's nozzle under oxygen supply conditions at the beginning and end of the blow cycle, and the average (T.Fe) in the slag when the carbon concentration in the molten steel is 0.04–0.05% by mass at the end of the blow cycle. The values ​​in parentheses in the column for the main gas back pressure (initial) in Table 7 are values ​​without control gas supply.

[0137] [Table 7]

[0138]

[0139] According to the results in Table 7, it can be seen that the amount of furnace dust generated is reduced when using spray guns G and H compared to spray gun F. However, the increase in iron oxide concentration in the slag reduces the effect of improving the iron yield. In addition, when using spray guns G and H, pre-desiliconization treatment of molten iron is sometimes required. However, this is not preferred because the decomposition of iron oxide contained in the desiliconizing agent generates endothermic heat.

[0140] In contrast, in this invention, even without pretreatment of the molten iron, slag ejection can be prevented by increasing the velocity of the top-blown oxygen jet when necessary by supplying control gas. Thus, without increasing the velocity of the top-blown oxygen jet, the jet velocity can be reduced to suppress furnace dust, and the increase in iron oxide concentration in the slag can be suppressed by supplying control gas at the end of refining, thereby enabling a stable and continuous improvement in iron yield. Furthermore, the reduction in iron oxide concentration in the slag during the above-described operation also has the advantage of saving on alloy iron used for deoxidation. In the cases of lances L and M, a slight increase in iron oxide concentration in the slag is preferable compared to other embodiments of the invention, thus reducing the improvement in iron yield. However, compared to conventional operations using lance F, the reduction in furnace dust generation and the improvement in iron yield are significant.

[0141] Industrial applications

[0142] It should be noted that the above embodiments describe the decarburization blowing process, but the present invention is not limited thereto, and the lance can also be used in dephosphorization blowing and desiliconization blowing. Furthermore, this technology can be applied even in refining processes using electric furnaces, as long as the refining process utilizes an oxygen-feeding lance. In particular, it is effective when it is desirable to increase the jet velocity or dynamic pressure without relying on changes in other gas supply conditions. For example, a refining method can be illustrated as follows: In the preliminary dephosphorization treatment of molten iron using a converter-type refining furnace, when the top-blown oxygen supply rate is reduced due to the decrease in dephosphorization oxygen efficiency at the end of refining, the oxygen-feeding refining method of the present invention, which uses a control gas to suppress the decrease in the top-blown jet velocity, suppresses the decrease in dephosphorization reaction efficiency.

Claims

1. A method for oxygen-assisted refining of molten iron, characterized in that, Oxygen-containing gas is blown into the molten iron through a top-blown lance to perform oxygen refining on the molten iron. During at least a portion of the oxygen refining process, oxygen-containing gas, as the main supply gas, is injected from the inlet side of the injection nozzle through the housing of the top-blown spray gun, while control gas is ejected from the outlet toward the nozzle. The outlet is located on the nozzle side where the nozzle's cross-sectional area is at its minimum in the nozzle's axial direction, or near the outlet, and is configured such that at least a portion of the outlet exists in both spaces when bisected by any plane passing through the nozzle's central axis. The spray outlet is provided in a slit shape along the entire circumference of the side of the spray nozzle. The axial length of the spray outlet from the spray nozzle is more than 0.1 times and less than 0.25 times the inner diameter of the nozzle at the part with the smallest cross-sectional area of ​​the spray nozzle. The interval between the slits is 0.6 mm to 2.0 mm. The pressure of the main supply gas at the inlet side of the injection nozzle is greater than the appropriate expansion pressure that satisfies the following equation (1): Yes / At=(5 5／2 / 6 3 )×(Fri / Night) －5／7 ×[1-(Fri / Night) 2／7 ] －1／2 ···(1) Here, At: the minimum cross-sectional area of ​​the injection nozzle, in mm. 2 Ae: Exit cross-sectional area of ​​the injection nozzle, in mm. 2 Pe: Ambient air pressure at the nozzle outlet, in kPa; Po: Appropriate expansion pressure of the nozzle, in kPa.

2. The oxygen-feeding refining method for molten iron according to claim 1, characterized in that, The portion of the nozzle whose cross-sectional area is the smallest in the nozzle axial direction is a portion where the cross-sectional area of ​​the nozzle is less than 1.1 times the smallest cross-sectional area in the nozzle axial direction.

3. The oxygen-feeding refining method for molten iron according to claim 1 or 2, characterized in that, As the injection nozzle, a straight pipe nozzle with a straight section having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, or a Laval nozzle with an end enlargement in the throat having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, is used.

4. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 3, characterized in that, The nozzle outlets are arranged circumferentially on the side of the injection nozzle in multiple directions. The product of the diameter of the inlet hole for introducing the control gas into the nozzle outlets and the number n of nozzle outlets for each injection nozzle is at least 0.4 times the nozzle inner diameter of the part with the smallest cross-sectional area of ​​the injection nozzle.

5. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 4, characterized in that, During at least a portion of the oxygen refining process, the flow rate of the control gas ejected toward the injection nozzle is at least 5% of the combined flow rate of the control gas and the main supply gas supplied to the injection nozzle.

6. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 5, characterized in that, The supply rate of the control gas is adjusted according to the supply rate of the oxygen-containing gas blown from the top-blown lance to the molten iron.

7. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 6, characterized in that, As the oxygen refining of the molten iron proceeds, the supply rate of the control gas is changed.

8. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 7, characterized in that, The supply rate of the control gas is adjusted according to the silicon concentration of the molten iron before the oxygen refining begins.

9. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 8, characterized in that, At the end of the oxygen refining process, after 85% of the total oxygen content in the oxygen-containing gas supplied in the oxygen refining process has been reached, the control gas is injected into the injection nozzle while the oxygen-containing gas, which is the main supply gas, is supplied.

10. The oxygen-feeding refining method for molten iron according to any one of claims 1 to 9, characterized in that, For molten iron with a silicon concentration of 0.40% or more before the start of oxygen refining, in the initial stage of oxygen refining before supplying 20% ​​of the total oxygen content in the oxygen-containing gas supplied during oxygen refining, the control gas is sprayed into the injection nozzle while the oxygen-containing gas, which is the main supply gas, is supplied.

11. A top-blowing spray gun, characterized in that, Used to blow oxygen-containing gas into the molten iron contained in the reaction vessel. The oxygen-containing gas injection nozzle, which penetrates the housing of the top-blowing spray gun, includes an outlet for ejecting control gas into the nozzle. This outlet is located on the nozzle side where the nozzle's cross-sectional area is at its minimum along the nozzle's axial direction, or near such a location. Furthermore, the nozzle is configured such that, when bisected by any plane passing through the nozzle's central axis, at least a portion of the outlet exists in both spaces. The spray outlet is provided in a slit shape along the entire circumference of the side of the spray nozzle. The axial length of the spray outlet from the spray nozzle is more than 0.1 times and less than 0.25 times the inner diameter of the nozzle at the part with the smallest cross-sectional area of ​​the spray nozzle. The interval between the slits is 0.6 mm to 2.0 mm. The multiple outlets for the control gas, which are arranged circumferentially along the sides of the nozzle in multiple directions, are interconnected within the top-blowing spray gun.

12. The top-blowing spray gun according to claim 11, characterized in that, The area near the part where the cross-sectional area of ​​the nozzle is the smallest in the nozzle axial direction is a part where the cross-sectional area of ​​the nozzle is less than 1.1 times the smallest cross-sectional area in the nozzle axial direction.

13. The top-blowing spray gun according to claim 11 or 12, characterized in that, The nozzle outlets are arranged circumferentially on the side of the injection nozzle in multiple directions. The product of the inner diameter of the nozzle from which the control gas is discharged, which is connected to the nozzle outlets, and the number n of nozzle outlets of each injection nozzle is at least 0.4 times the inner diameter of the nozzle corresponding to the minimum cross-sectional area of ​​the injection nozzle.

14. A top-blowing spray gun, characterized in that, Used to blow oxygen-containing gas into the molten iron contained in the reaction vessel. The oxygen-containing gas injection nozzle, which penetrates the housing of the top-blowing spray gun, has an outlet for ejecting control gas toward the nozzle. This outlet is located circumferentially on the nozzle side at or near a location where the cross-sectional area is smallest in the nozzle axial direction, and is provided in a slit-like manner throughout the entire circumference. The axial length of the spray nozzle at the spray outlet is more than 0.1 times and less than 0.25 times the inner diameter of the nozzle at the part with the smallest cross-sectional area of ​​the spray nozzle, and the interval of the slit gap is 0.6 mm to 2.0 mm.

15. The top-blowing spray gun according to claim 14, characterized in that, The area near the part where the cross-sectional area of ​​the nozzle is the smallest in the nozzle axial direction is a part where the cross-sectional area of ​​the nozzle is less than 1.1 times the smallest cross-sectional area in the nozzle axial direction.

16. The top-blowing spray gun according to any one of claims 11 to 15, characterized in that, As the injection nozzle, a straight pipe nozzle with a straight section having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, or a Laval nozzle with an end enlargement in the throat having the smallest cross-sectional area in the nozzle axial direction adjacent to the nozzle outlet, is used.

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