A high-purity duplex stainless steel alloy and a preparation process thereof

Abstract

The application relates to the technical field of metallurgy and special steel refining, and discloses a high-purity duplex stainless steel alloy and a preparation process thereof, wherein the alloy contains C 0.015%-0.030%, Cr 22.0%-23.0%, Ni 4.5%-6.5%, Mo 3.0%-3.5%, N 0.14%-0.20%, Si 0.3%-1.0%, Mn 0.8%-2.0%, P<=0.02%, and the balance of Fe and inevitable impurities. In the refining process, after pre-deoxidation, silicon-barium alloy wires are first fed, and then rare earth cerium-iron alloy wires and nickel-magnesium slow-release alloy wires are synchronously fed, so that the inclusion modification and in-situ micro-bubble floatation synergistic effect are realized. The method is beneficial to shortening the inclusion removal time, reducing the temperature drop, improving the nitrogen preservation effect, and reducing the total oxygen and total sulfur contents in the final casting blank.

Description

Technical Field

[0001] This invention relates to the field of metallurgy and special steel refining technology, and in particular to a high-purity duplex stainless steel alloy and its preparation process. Background Technology

[0002] Duplex stainless steel is widely used in petrochemical, marine engineering, and pressure vessel industries due to its high strength, good corrosion resistance, and good weldability. This type of steel has a high content of elements such as chromium, molybdenum, and nitrogen, which places high demands on the purity of the molten steel and the control of inclusions. In actual production, calcium treatment is usually used in the LF refining stage for deoxidation and inclusion modification. However, due to the high reactivity and low boiling point of calcium, in high-alloy duplex stainless steel systems, unstable reactions and insufficient inclusion modification can easily occur, leading to the formation of high-melting-point solid inclusions, which affect the aggregation and flotation removal of inclusions.

[0003] Meanwhile, existing processes often rely on prolonged weak argon blowing to promote the flotation and removal of inclusions. However, this method can easily lead to a large temperature drop in the molten steel, increasing the difficulty of subsequent casting control. Furthermore, it can exacerbate nitrogen escape, which is detrimental to the stable control of nitrogen content in duplex stainless steel. For duplex stainless steel that requires a balance of high purity, low oxygen and sulfur content, and nitrogen retention, existing refining methods still suffer from insufficient inclusion removal efficiency, low process stability, and large fluctuations in finished product purity. Therefore, it is necessary to propose a new refining control scheme. Summary of the Invention

[0004] The technical problem to be solved by this invention is that, in the existing duplex stainless steel refining process, when using single calcium treatment for deoxidation and inclusion modification, high-melting-point solid magnesium aluminum spinel or hard sulfur oxide inclusions are easily generated. At the same time, the subsequent long-term weak argon blowing flotation operation will cause the steel liquid to cool down and nitrogen to escape, resulting in the problem that inclusions are difficult to float and remove effectively, the process stability is poor, and the purity and nitrogen retention of the final steel liquid are insufficient.

[0005] To solve the above-mentioned technical problems, the present invention provides the following technical solution.

[0006] In a first aspect, the present invention provides a high-purity duplex stainless steel alloy, the alloy being composed of the following components by mass percentage: C: 0.015%–0.030%, Cr: 22.0%–23.0%, Ni: 4.5%–6.5%, Mo: 3.0%–3.5%, N: 0.14%–0.20%, Si: 0.3%–1.0%, Mn: 0.8%–2.0%, P ≤ 0.02%, with the balance being Fe and unavoidable impurities;

[0007] Furthermore, the alloy is prepared by a process including the following refining process: after refining and pre-deoxidation, silicon-barium alloy wire is fed into the deep part of the molten steel in stages, and rare earth cerium-iron alloy wire and nickel-magnesium slow-release alloy wire are fed in simultaneously; the silicon-barium alloy wire is used to construct a low-melting-point barium-based composite liquid phase in the deep part of the molten steel to absorb and modify inclusions, and the nickel-magnesium slow-release alloy wire is used to gently release in-situ magnesium microbubbles when heated, promoting the flotation and removal of non-metallic inclusions after rare earth modification treatment.

[0008] By adopting the above technical solution, this invention mainly utilizes the coupled control principle of thermodynamic phase change and kinetic air flotation mass transfer to achieve the following technical effects:

[0009] Specifically, after refining and pre-deoxidation, a barium-silicon alloy wire is fed into the deep part of the molten steel. Since barium has a higher boiling point than the temperature of the molten steel, it usually participates in the reaction in the molten steel in a liquid state, and preferentially reacts with inclusions such as alumina and silicon dioxide in the molten steel to generate dispersed low-melting-point barium-aluminum-silicon composite liquid droplets.

[0010] Based on this, when rare earth cerium-iron alloy wire and nickel-magnesium slow-release alloy wire are subsequently fed in simultaneously, the physicochemical conditions of the entire molten steel system change. Normally, magnesium exists in the nickel matrix in a Ni-Mg bound state, which to some extent inhibits the thermodynamic activity of magnesium in high-temperature molten steel. When the nickel-magnesium slow-release alloy wire enters the molten steel at 1590℃~1610℃, the nickel matrix produces a slow-release effect, transforming the instantaneous vaporization of magnesium into a relatively controlled steady-state micro-vaporization, thereby generating micron-sized magnesium vapor bubbles in situ deep within the molten steel.

[0011] Simultaneously, with the highly reactive cerium and magnesium entering the molten steel, the nascent cerium-oxygen-sulfur reactants can directly use the pre-existing barium-based composite liquid droplets as a heterogeneous nucleation substrate. Due to the fluxing effect of the barium-based droplets, the surface tension and activation energy of the relevant inclusion nucleation interface are reduced, which is conducive to their formation and evolution in a state of lower crystallization and lower interface energy.

[0012] Subsequently, the modified composite inclusions come into contact with micron-sized magnesium vapor bubbles generated by the slow release of nickel and magnesium in the same space. Based on the principle of minimizing surface tension, the magnesium vapor bubbles preferentially adhere to the surface of the non-metallic inclusions, thereby facilitating the formation of a three-phase interfacial adsorption state consisting of gaseous magnesium vapor, a liquid barium-based flux layer, and a solid cerium metamorphic nucleus. The magnesium vapor bubbles provide additional buoyancy to the inclusions, changing the previous floating mode that relied solely on gravity difference, and promoting the floating and discharge of the composite inclusions.

[0013] The above reaction process largely avoids strong interfacial turbulence and secondary oxygen absorption during refining, reducing non-metallic hard pitting corrosion sources. Microbubble kinetic flotation helps shorten the weak stirring flotation cycle and alleviates the temperature drop of molten steel caused by prolonged argon blowing, thus promoting the stable retention of nitrogen in the alloy system and providing favorable conditions for obtaining a stable dual-phase microstructure. Furthermore, by reducing micro-crack initiation sources such as incompletely modified spinel and sulfur oxide aggregates, the critical pitting temperature and resistance to localized corrosion of the alloy can be improved to some extent.

[0014] Preferably, the silicon-barium alloy wire is made of a powdered core material wrapped in an outer layer of low-carbon steel strip, the powdered core material containing 10.0% to 15.0% barium and 40.0% to 50.0% silicon by weight; the rare earth cerium-iron alloy wire is made of a powdered core material wrapped in an outer layer of low-carbon steel strip, the powdered core material containing 25.0% to 30.0% cerium by weight; the nickel-magnesium slow-release alloy wire is made of nickel-magnesium alloy powder coated in an outer layer of cold-rolled low-carbon steel strip, the particle size of the nickel-magnesium alloy powder is 75μm to 150μm, and the nickel-magnesium alloy powder is composed of the following components by weight percentage: magnesium 4.0% to 8.0%, the balance being nickel and unavoidable impurities.

[0015] By adopting the above technical solutions, the specific composition and structure of each alloy wire are defined. The specific ratio of silicon-barium alloy wire ensures the formation of sufficient liquid barium-based composite microdroplets in the molten steel; the specific ratio of rare earth cerium-iron alloy wire provides a suitable cerium source for inclusion modification; the nickel-magnesium alloy powder with specific particle size and composition is conducive to the existence of magnesium elements in a bound state under the constraint of the nickel matrix, thereby achieving a stable and gradual release of magnesium bubbles during the wire feeding process and minimizing the violent turbulence of the molten steel caused by the instantaneous vaporization of pure magnesium.

[0016] Preferably, the nickel-magnesium alloy powder is prepared by the following method: the amount of raw materials is calculated according to the target ratio, and an additional 20.0% of the target magnesium mass of pure magnesium blocks is added as burn-off compensation; the pure nickel blocks are placed in a vacuum induction melting furnace and heated to 1500°C to melt pure nickel liquid; high-purity argon gas is introduced into the melting furnace, and the pure magnesium blocks are pressed into the bottom of the pure nickel liquid using a perforated refractory bell jar, while maintaining induction electromagnetic stirring to alloy it; the obtained alloy liquid is cast and cooled, crushed, ball-milled for 120 minutes under argon protection, and sieved.

[0017] By employing the above technical solution, combining the bell-jar pressing method with induction electromagnetic stirring, low-melting-point and low-boiling-point pure magnesium can be effectively dissolved in 1500℃ pure nickel liquid. Increasing the amount of pure magnesium compensation offsets the volatilization loss during the high-temperature smelting process, helping to ensure the accuracy of the magnesium content in the final alloy powder. Argon-protected ball milling reduces surface oxidation of the alloy powder during the crushing and refining process, which is beneficial for maintaining the purity of the nickel-magnesium slow-release alloy wire core material.

[0018] Preferably, the alloy has a total oxygen content of 6.5 ppm to 7.8 ppm and a total sulfur content of 3.5 ppm to 4.6 ppm.

[0019] By adopting the above technical solution, through the front-end barium-cerium composite modification and in-situ magnesium bubble flotation process, the flotation removal rate of inclusions inside the molten steel is improved, and the characteristic peaks of the relevant highly crystalline magnesium-aluminum spinel phase and rare earth oxide phase are significantly weakened, thereby enabling the alloy matrix to achieve a lower oxygen and sulfur content level and possess better physical purity.

[0020] Secondly, the present invention also provides a process for preparing a high-purity duplex stainless steel alloy, comprising the following steps:

[0021] S1. The primary steel molten steel is poured into the AOD furnace, bottom-blown mixed gas is introduced to carry out decarburization reaction, alloy is added to carry out alloying, the tapping temperature is controlled and the steel molten steel is poured into the LF refining ladle, and strong deoxidation is not carried out during the tapping process.

[0022] S2. After the ladle is placed in the LF furnace, the electrodes are lowered for heating, slag-forming material is added for slag formation, aluminum-manganese-iron alloy is added for pre-deoxidation, and at the same time, bottom blowing argon gas is turned on in the ladle to form a strong stirring state and is maintained.

[0023] S3. While maintaining strong stirring with bottom-blown argon gas, feed the silicon-barium alloy wire into the deep part of the molten steel, and continue to maintain strong stirring after the wire feeding is completed.

[0024] S4. Reduce the bottom blowing argon flow rate to a weak stirring state, and then start the dual-line parallel feeding system to simultaneously feed the rare earth cerium iron alloy wire and nickel magnesium slow-release alloy wire into the deep part of the molten steel, so that the two alloy wires melt and release simultaneously on the same horizontal plane.

[0025] S5. After the dual-line feeding is completed, maintain a weak stirring state to perform air flotation and impurity removal. Then, further reduce the bottom blowing argon gas to a soft blowing state and maintain it, and then turn off the bottom blowing argon gas.

[0026] S6. The refined steel ladle is transported to the continuous casting turret and continuously cast using a long nozzle argon-protected casting process to obtain duplex stainless steel billets.

[0027] By adopting the above technical solution, this process relies on the switching of the strength of the bottom-blowing flow field and the corresponding alloy wire feeding sequence to achieve coordinated control of the refining process, specifically achieving the following technical effects:

[0028] During the initial pre-deoxidation stage of refining, bottom-blown argon gas is used to keep the molten steel under strong stirring. The inclusions generated during pre-deoxidation collide and aggregate under the convection of the molten steel. With the help of this strong stirring condition, the silicon-barium alloy wire is fed in, and the barium element released by melting can diffuse quickly in the entire molten steel, thereby creating a more uniform barium-based liquid phase modification environment for subsequent reactions.

[0029] After entering the bottom-blowing and weak stirring stage, the shear stress of the fluid inside the molten steel decreases accordingly. At this time, rare earth cerium-iron alloy wire and nickel-magnesium slow-release alloy wire are simultaneously fed in. This operation arrangement places the rare earth modification reaction and magnesium bubble release in a relatively gentle flow field, which is conducive to prolonging the contact time between magnesium vapor bubbles and non-metallic inclusions, and better utilizing the flotation effect of in-situ microbubbles to lift the modified composite inclusions to the steel-slag interface. At the same time, since the cycle of impurity removal relying solely on long-term weak argon blowing in conventional processes is shortened, it is also beneficial to control excessive cooling of the molten steel and nitrogen escape.

[0030] Preferably, in step S1, a decarburization reaction is carried out until the C content in the molten steel drops to 0.015%–0.030%; the bottom-blowing gas is adjusted to a mixture of nitrogen and argon to alloy the N content to 0.18%–0.20%; the tapping temperature is controlled at 1620℃–1640℃; in step S2, heating is controlled to stabilize the molten steel temperature at 1590℃–1610℃; the mass ratio of calcium oxide to silicon dioxide in the refining top slag is controlled at 3.5–4.0; the bottom-blowing argon flow rate under strong stirring is 35NL / min–45NL / min, and the holding time is 5–8 minutes.

[0031] By adopting the above technical solutions, controlling the tapping and refining temperatures within a specific range helps ensure the reaction conditions during the alloying process. Maintaining the basicity of the refining top slag between 3.5 and 4.0 typically allows for better absorption and fixation of inclusions floating to the slag-metal interface, reducing the risk of impurities returning to the molten steel. The specific flow rate and holding time of the bottom-blowing strong stirring promote uniform element diffusion and provide conditions for subsequent barium-based liquid phase modification.

[0032] Preferably, in step S3, the feeding speed of the silicon-barium alloy wire is 2.5 m / s to 3.0 m / s, and the amount of barium added, calculated as pure barium, is 0.03% to 0.05% of the total mass of the molten steel; the strong stirring time after the wire feeding is completed is 3 to 5 minutes; in step S4, the synchronous feeding operation is carried out within 30 to 60 seconds after the weak stirring is started; the downward speed of the two alloy wires is adjusted to 1.5 m / s to 2.5 m / s; wherein, calculated as pure cerium, the amount of rare earth cerium-iron alloy wire added is 0.01% to 0.02% of the total mass of the molten steel; and calculated as pure magnesium, the amount of nickel-magnesium slow-release alloy wire added is 0.003% to 0.008% of the total mass of the molten steel.

[0033] By adopting the above technical solution and setting a relatively fast wire feeding speed, the main purpose is to enable the alloy wire to penetrate the slag layer above the molten steel and enter the depths of the ladle, preventing the active alloying elements from melting prematurely on the shallow surface and causing oxidation and burn-off. Simultaneous feeding of both wires within 30-60 seconds after the start of weak stirring is to allow buffer time for the molten steel flow field to transition from a highly turbulent state to a relatively stable state. Furthermore, controlling the amount of effective elements added to each alloy wire within a specified range can avoid excessive disturbance to the molten steel system while meeting the requirements for inclusion modification and air flotation impurity removal.

[0034] Preferably, in step S5, the time for maintaining the weak stirring state is 5 to 8 minutes, and the bottom blowing argon flow rate in the weak stirring state is 10 NL / min to 15 NL / min; the holding time of the soft blowing state is 5 minutes; in step S6, the superheat of the molten steel in the continuous casting tundish is controlled at 20°C to 25°C.

[0035] By adopting the above technical solution, the flow rate and duration of the weak stirring are matched with the rising process of in-situ magnesium bubbles in the molten steel, which helps to allow sufficient time for inclusions to float into the slag layer. The soft blowing treatment after the weak stirring usually allows the remaining fine inclusions to continue to collide and aggregate, while also contributing to further homogenization of the composition and temperature of the entire ladle of molten steel. Controlling the appropriate superheat in the tundish during continuous casting helps to improve the stability of the casting process and ensure the quality of the cast billet.

[0036] Preferably, the nickel-magnesium slow-release alloy wire used in step S4 is prepared in advance by the following process: the amount of raw materials is calculated according to the target alloy mass fraction of 4.0% to 8.0% magnesium and 92.0% to 96.0% nickel, and an additional 20.0% pure magnesium block is added as burn-off compensation; the pure nickel block is placed in the crucible of a vacuum induction melting furnace, the vacuum is drawn to below 10 Pa, and then the temperature is raised to 1500°C to completely melt the pure nickel block to obtain pure nickel liquid; high-purity argon gas of 0.8 MPa is introduced into the melting furnace, and a perforated slurry is used to melt the slurry. A refractory bell jar is used to quickly press pure magnesium blocks into the bottom of pure nickel liquid, and induction electromagnetic stirring is maintained for 5 minutes to fully alloy them. The resulting alloy liquid is cast into a water-cooled copper mold for cooling to obtain a nickel-magnesium alloy ingot. After coarse crushing, it is placed in a ball mill and ground for 120 minutes under argon protection. The nickel-magnesium alloy powder with a particle size of 75μm to 150μm is collected by sieving. The cold-rolled low-carbon steel strip is bent into a U-shaped groove by the unit, the nickel-magnesium alloy powder is added, closed and compacted, and drawn into a cored wire with a wire diameter of 13.0mm, with the powder filling amount controlled at 150g / m.

[0037] By adopting the above technical solution, a preparation process for nickel-magnesium slow-release alloy wire was specifically designed. Since magnesium has a low boiling point, pressing it into the high-temperature nickel melt using a perforated refractory bell jar is beneficial for ensuring that magnesium exists in a bound state constrained by the nickel matrix. Subsequently, crushing, grinding, and core-drawing are performed under argon protection. This is mainly because magnesium-rich alloy powder is prone to oxidation in air, thus reducing the oxidation risk during processing and maintaining the slow-release effect of the cored wire during subsequent ladle feeding.

[0038] In summary, the present invention has at least one of the following beneficial technical effects:

[0039] 1. This invention establishes a feeding sequence that matches the strength of the bottom blowing flow field by first feeding a silicon-barium alloy wire after pre-deoxidation, followed by the simultaneous feeding of a rare earth cerium-iron alloy wire and a nickel-magnesium slow-release alloy wire. This facilitates the synergistic implementation of barium-based liquid phase modification, rare earth alteration, and in-situ microbubble flotation in the deep part of the molten steel.

[0040] 2. By introducing a silicon-barium alloy wire during the strong stirring stage, this invention facilitates the rapid diffusion of barium elements in the fully encased molten steel and the construction of a more uniform barium-based composite liquid phase environment, providing a foundation for subsequent inclusion modification and flotation removal.

[0041] 3. This invention simultaneously feeds rare earth cerium-iron alloy wire and nickel-magnesium slow-release alloy wire during the weak stirring stage, so that the rare earth modification reaction and magnesium vapor bubble release are in a relatively gentle flow field environment, which is conducive to prolonging the contact time between magnesium bubbles and inclusions, thereby improving the efficiency of inclusion flotation and removal.

[0042] 4. By using a nickel-magnesium slow-release alloy wire instead of the conventional method of adding pure magnesium, this invention helps to reduce the risk of violent interfacial turbulence and abnormal reactions caused by concentrated magnesium gasification, and improves the stability of the refining process.

[0043] 5. This invention improves nitrogen retention by shortening the time required for impurity removal in conventional processes that rely on prolonged weak argon blowing. This helps reduce the temperature drop of molten steel and suppress nitrogen escape. Attached Figure Description

[0044] Figure 1 Thermogravimetric-differential scanning calorimetry (TG) curves of pure magnesium powder and nickel-magnesium alloy powder obtained in Preparation Example 2 are shown below. (a) is the TG curve, and (b) is the DSC curve.

[0045] Figure 2 The X-ray diffraction patterns of inclusions extracted by electrolysis in Example 1 and Comparative Example 4 of this invention are shown. Detailed Implementation

[0046] The main raw materials and reagents used in the following examples and comparative examples are as follows. Unless otherwise specified, all raw materials, excipients and reagents are commercially available industrial grade, analytical grade or higher grade products.

[0047] The silicon-barium alloy wire is made of a powdered core material wrapped in an outer layer of low-carbon steel strip. The wire diameter is 13.0 mm and the coating thickness is 0.4 mm. The powdered core material contains 10.0% to 15.0% barium, 40.0% to 50.0% silicon, ≤1.5% aluminum, ≤0.5% carbon, and the balance is iron.

[0048] Rare earth cerium-iron alloy wire is made of powdered core material wrapped in an outer layer of low carbon steel strip. The wire diameter is 13.0 mm and the coating thickness is 0.4 mm. The mass fraction of cerium in the powdered core material is 25.0% to 30.0%, the mass fraction of silicon is ≤5.0%, the mass fraction of manganese is ≤2.0%, and the balance is iron.

[0049] An aluminum-manganese-iron alloy, wherein the mass fraction of aluminum is 15.0%–20.0%, the mass fraction of manganese is 35.0%–40.0%, the mass fraction of carbon is ≤2.0%, the mass fraction of silicon is ≤1.5%, the mass fraction of phosphorus is ≤0.05%, the mass fraction of sulfur is ≤0.04%, and the balance is iron.

[0050] Cold-rolled low-carbon steel strip with a carbon mass fraction of ≤0.08%, a width of 50.0 mm, and a thickness of 0.4 mm.

[0051] Preparation Example 1:

[0052] This preparation example provides a method for preparing a nickel-magnesium slow-release alloy wire, including the following steps:

[0053] (1) Calculate the amount of raw materials according to the target alloy mass fraction of 4.0% magnesium and 96.0% nickel, and add an extra 20.0% pure magnesium block as burn loss compensation for later use;

[0054] (2) Place the pure nickel block in the alumina crucible of the vacuum induction melting furnace, evacuate to below 10 Pa, and then heat to 1500℃ to completely melt the pure nickel block to obtain pure nickel liquid;

[0055] (3) Introduce 0.8MPa of high-purity argon into the melting furnace, place the pure magnesium block in the perforated refractory bell jar, use the bell jar to quickly press the pure magnesium block into the bottom of the pure nickel liquid, maintain induction electromagnetic stirring for 5 minutes to fully alloy it.

[0056] (4) The obtained alloy liquid is quickly poured into a water-cooled copper mold and cooled to room temperature to obtain a nickel-magnesium alloy ingot;

[0057] (5) The nickel-magnesium alloy ingot is coarsely crushed by a jaw crusher, and then ground in a ball mill for 120 min under argon protection. The nickel-magnesium alloy powder with a particle size of 75μm to 150μm is collected by sieving.

[0058] (6) The cold-rolled low-carbon steel strip is bent into a U-shaped groove by the cored wire unit. Nickel-magnesium alloy powder is continuously and evenly added into the U-shaped groove, then closed and compacted, and drawn into a cored wire with a diameter of 13.0 mm. The powder filling amount is controlled to be 150 g / m to obtain a nickel-magnesium slow-release alloy wire.

[0059] (7) The chemical composition of the nickel-magnesium alloy powder obtained in step (5) was analyzed. The results showed that the nickel-magnesium alloy powder contained 4.1% magnesium and 95.9% nickel by mass percentage.

[0060] Preparation Example 2:

[0061] This preparation example provides a method for preparing a nickel-magnesium slow-release alloy wire, including the following steps:

[0062] (1) Calculate the amount of raw materials according to the target alloy mass fraction of 6.0% magnesium and 94.0% nickel, and add an extra 20.0% pure magnesium block as burn loss compensation for later use;

[0063] (2) Place the pure nickel block in the alumina crucible of the vacuum induction melting furnace, evacuate to below 10 Pa, and then heat to 1500℃ to completely melt the pure nickel block to obtain pure nickel liquid;

[0064] (3) Introduce 0.8MPa of high-purity argon into the melting furnace, place the pure magnesium block in the perforated refractory bell jar, use the bell jar to quickly press the pure magnesium block into the bottom of the pure nickel liquid, maintain induction electromagnetic stirring for 5 minutes to fully alloy it.

[0065] (4) The obtained alloy liquid is quickly poured into a water-cooled copper mold and cooled to room temperature to obtain a nickel-magnesium alloy ingot;

[0066] (5) The nickel-magnesium alloy ingot is coarsely crushed by a jaw crusher, and then ground in a ball mill for 120 min under argon protection. The nickel-magnesium alloy powder with a particle size of 75μm to 150μm is collected by sieving.

[0067] (6) The cold-rolled low-carbon steel strip is bent into a U-shaped groove by the cored wire unit. Nickel-magnesium alloy powder is continuously and evenly added into the U-shaped groove, then closed and compacted, and drawn into a cored wire with a diameter of 13.0 mm. The powder filling amount is controlled to be 150 g / m to obtain a nickel-magnesium slow-release alloy wire.

[0068] (7) The chemical composition of the nickel-magnesium alloy powder obtained in step (5) was analyzed. The results showed that the nickel-magnesium alloy powder contained 6.1% magnesium and 93.9% nickel by mass percentage.

[0069] Preparation Example 3:

[0070] This preparation example provides a method for preparing a nickel-magnesium slow-release alloy wire, including the following steps:

[0071] (1) Calculate the amount of raw materials according to the target alloy mass fraction of 8.0% magnesium and 92.0% nickel, and add an extra 20.0% pure magnesium block as burn loss compensation for later use;

[0072] (2) Place the pure nickel block in the alumina crucible of the vacuum induction melting furnace, evacuate to below 10 Pa, and then heat to 1500℃ to completely melt the pure nickel block to obtain pure nickel liquid;

[0073] (3) Introduce 0.8MPa of high-purity argon into the melting furnace, place the pure magnesium block in the perforated refractory bell jar, use the bell jar to quickly press the pure magnesium block into the bottom of the pure nickel liquid, maintain induction electromagnetic stirring for 5 minutes to fully alloy it.

[0074] (4) The obtained alloy liquid is quickly poured into a water-cooled copper mold and cooled to room temperature to obtain a nickel-magnesium alloy ingot;

[0075] (5) The nickel-magnesium alloy ingot is coarsely crushed by a jaw crusher, and then ground in a ball mill for 120 min under argon protection. The nickel-magnesium alloy powder with a particle size of 75μm to 150μm is collected by sieving.

[0076] (6) The cold-rolled low-carbon steel strip is bent into a U-shaped groove by the cored wire unit. Nickel-magnesium alloy powder is continuously and evenly added into the U-shaped groove, then closed and compacted, and drawn into a cored wire with a diameter of 13.0 mm. The powder filling amount is controlled to be 150 g / m to obtain a nickel-magnesium slow-release alloy wire.

[0077] (7) The chemical composition of the nickel-magnesium alloy powder obtained in step (5) was analyzed. The results showed that the nickel-magnesium alloy powder contained 7.9% magnesium and 92.1% nickel by mass percentage.

[0078] Example 1:

[0079] This embodiment provides a preparation process for a high-purity duplex stainless steel alloy, including the following steps:

[0080] (1) The primary molten steel from the electric arc furnace is added to the AOD furnace, and a bottom-blown O2 / Ar mixed gas is introduced to carry out a decarburization reaction until the C content in the molten steel drops to 0.02%. High-carbon ferrochrome, ferromolybdenum, and pure nickel plates are added for alloying. The amount of pure nickel plates added is determined based on the target Ni content and the amount of Ni brought in by the subsequently fed nickel-magnesium slow-release alloying line. The bottom-blown gas is adjusted to an N2 / Ar mixed gas to alloy the N content in the molten steel to 0.19%. The tapping temperature is controlled at 1630℃, and the molten steel is added to the LF refining ladle. No strong deoxidation is carried out during the tapping process.

[0081] (2) After the ladle is positioned in the LF furnace, three-phase electrodes are lowered for heating to stabilize the molten steel temperature at 1600℃. Activated lime and fluorite are added for slag formation, and the binary basicity (CaO / SiO2 mass ratio) of the refining top slag is controlled at 3.8. After the top slag melts, aluminum-manganese-ferroalloy is added for pre-deoxidation, and at the same time, bottom blowing argon gas is turned on in the ladle, with the flow rate set to strong stirring (40 NL / min) and maintained for 6 minutes.

[0082] (3) Under the conditions of maintaining the molten steel temperature at 1600℃ and strong bottom-blown argon stirring (40NL / min), a silicon-barium alloy wire with a barium mass fraction of 12.5% ​​in the powdered core material is fed into the deep part of the molten steel through a wire feeder. The feeding speed is controlled at 2.8m / s, and the amount of barium added is 0.04% of the total mass of the molten steel, based on pure barium element. After the wire feeding is completed, strong stirring (40NL / min) is maintained for another 4 minutes.

[0083] (4) Reduce the bottom-blown argon flow rate to a weak stirring state (12 NL / min). Within 45 seconds of starting the weak stirring, start the dual-wire parallel feeding system and simultaneously feed rare earth cerium-iron alloy wire with a cerium mass fraction of 28.0% in the powdered core material and the nickel-magnesium slow-release alloy wire obtained in Preparation Example 2 into the deep part of the molten steel (400 mm from the bottom of the ladle). The amount of rare earth cerium-iron alloy wire added is 0.015% of the total mass of the molten steel based on pure cerium; the amount of nickel-magnesium slow-release alloy wire added is 0.005% of the total mass of the molten steel based on pure magnesium. Adjust the downward linear velocity of the two alloy wires to 2.0 m / s so that they melt and release synchronously on the same horizontal plane.

[0084] (5) After the dual-line feeding is completed, maintain a weak stirring state (12NL / min) for 6 minutes. After the weak stirring is completed, take a sample to confirm that the N content of the molten steel is within the target range. Further reduce the bottom blowing argon gas to a soft blowing state and maintain it for 5 minutes, then turn off the bottom blowing argon gas.

[0085] (6) The LF-refined steel ladle is transported to the continuous casting turret and cast using a long nozzle argon protection process. The superheat of the molten steel in the tundish is controlled at 22°C, and the steel is continuously cast to obtain duplex stainless steel billets.

[0086] (7) The chemical composition of the obtained duplex stainless steel billet was analyzed. The results showed that the composition of the billet by mass percentage was: C 0.020%, Cr 22.54%, Ni 5.72%, Mo 3.18%, N 0.186%, Si 0.56%, Mn 1.38%, P 0.017%, with the balance being Fe and unavoidable impurities.

[0087] Example 2:

[0088] This embodiment provides a preparation process for a high-purity duplex stainless steel alloy, including the following steps:

[0089] (1) The primary molten steel from the electric arc furnace is added to the AOD furnace, and a bottom-blown O2 / Ar mixed gas is introduced to carry out a decarburization reaction until the C content in the molten steel drops to 0.015%. High-carbon ferrochrome, ferromolybdenum, and pure nickel plates are added for alloying. The amount of pure nickel plates added is determined based on the target Ni content and the amount of Ni brought in by the subsequently fed nickel-magnesium slow-release alloying line. The bottom-blown gas is adjusted to an N2 / Ar mixed gas to alloy the N content in the molten steel to 0.18%. The tapping temperature is controlled at 1620℃, and the molten steel is added to the LF refining ladle. No strong deoxidation is carried out during the tapping process.

[0090] (2) After the ladle is positioned in the LF furnace, three-phase electrodes are lowered for heating to stabilize the molten steel temperature at 1590℃. Activated lime and fluorite are added for slag formation, and the binary basicity (CaO / SiO2 mass ratio) of the refining top slag is controlled at 3.5. After the top slag melts, aluminum-manganese-ferroalloy is added for pre-deoxidation, and at the same time, bottom blowing argon gas is turned on in the ladle, with the flow rate set to strong stirring (35 NL / min), and maintained for 5 minutes.

[0091] (3) Under the conditions of maintaining the molten steel temperature at 1590℃ and strong bottom-blown argon stirring (35NL / min), a silicon-barium alloy wire with a barium mass fraction of 10.0% in the powdered core material is fed into the deep part of the molten steel through a wire feeder. The feeding speed is controlled at 2.5m / s, and the amount of barium added is 0.03% of the total mass of the molten steel, based on pure barium element. After the wire feeding is completed, strong stirring (35NL / min) is maintained for 3 minutes.

[0092] (4) Reduce the bottom-blown argon flow rate to a weak stirring state (10 NL / min). Within 60 seconds of starting the weak stirring, start the dual-wire parallel feeding system and simultaneously feed rare earth cerium-iron alloy wire with a cerium mass fraction of 25.0% in the powdered core material and the nickel-magnesium slow-release alloy wire obtained in Preparation Example 1 into the deep part of the molten steel (300 mm from the bottom of the ladle). The amount of rare earth cerium-iron alloy wire added is 0.01% of the total mass of the molten steel based on pure cerium; the amount of nickel-magnesium slow-release alloy wire added is 0.003% of the total mass of the molten steel based on pure magnesium. Adjust the downward linear velocity of the two alloy wires to 1.5 m / s so that they melt and release synchronously on the same horizontal plane.

[0093] (5) After the dual-line feeding is completed, maintain a weak stirring state (10NL / min) for 5 minutes. After the weak stirring is completed, take a sample to confirm that the N content of the molten steel is within the target range. Further reduce the bottom blowing argon gas to a soft blowing state and maintain it for 5 minutes, then turn off the bottom blowing argon gas.

[0094] (6) The LF-refined steel ladle is transported to the continuous casting turret and cast using a long nozzle argon protection process. The superheat of the molten steel in the tundish is controlled at 20°C, and the steel is continuously cast to obtain duplex stainless steel billets.

[0095] (7) The chemical composition of the obtained duplex stainless steel billet was analyzed. The results showed that the composition of the billet by mass percentage was: C 0.015%, Cr 22.12%, Ni 5.06%, Mo 3.05%, N 0.177%, Si 0.41%, Mn 1.05%, P 0.014%, with the balance being Fe and unavoidable impurities.

[0096] Example 3:

[0097] This embodiment provides a preparation process for a high-purity duplex stainless steel alloy, including the following steps:

[0098] (1) The primary molten steel from the electric arc furnace is added to the AOD furnace, and a bottom-blown O2 / Ar mixed gas is introduced to carry out a decarburization reaction until the C content in the molten steel drops to 0.03%. High-carbon ferrochrome, ferromolybdenum, and pure nickel plates are added for alloying. The amount of pure nickel plates added is determined based on the target Ni content and the amount of Ni brought in by the subsequently fed nickel-magnesium slow-release alloying line. The bottom-blown gas is adjusted to an N2 / Ar mixed gas to alloy the N content in the molten steel to 0.20%. The tapping temperature is controlled at 1640℃, and the molten steel is added to the LF refining ladle. No strong deoxidation is performed during the tapping process.

[0099] (2) After the ladle is positioned in the LF furnace, three-phase electrodes are lowered for heating to stabilize the molten steel temperature at 1610℃. Activated lime and fluorite are added for slag formation, and the binary basicity (CaO / SiO2 mass ratio) of the refining top slag is controlled at 4.0. After the top slag melts, aluminum-manganese-ferroalloy is added for pre-deoxidation, and at the same time, bottom blowing argon gas is turned on in the ladle, with the flow rate set to strong stirring (45 NL / min), and maintained for 8 minutes.

[0100] (3) Under the conditions of maintaining the molten steel temperature at 1610℃ and strong bottom-blown argon stirring (45NL / min), a silicon-barium alloy wire with a barium mass fraction of 15.0% in the powdered core material is fed into the deep part of the molten steel through a wire feeder. The feeding speed is controlled at 3.0m / s, and the amount of barium added is 0.05% of the total mass of the molten steel, based on pure barium element. After the wire feeding is completed, strong stirring (45NL / min) is maintained for 5 minutes.

[0101] (4) Reduce the bottom-blown argon flow rate to a weak stirring state (15 NL / min). Within 30 seconds of starting the weak stirring, start the dual-wire parallel feeding system and simultaneously feed rare earth cerium-iron alloy wire with a cerium mass fraction of 30.0% in the powdered core material and the nickel-magnesium slow-release alloy wire obtained in Preparation Example 3 into the deep part of the molten steel (500 mm from the bottom of the ladle). The amount of rare earth cerium-iron alloy wire added is 0.02% of the total mass of the molten steel based on pure cerium; the amount of nickel-magnesium slow-release alloy wire added is 0.008% of the total mass of the molten steel based on pure magnesium. Adjust the downward linear velocity of the two alloy wires to 2.5 m / s so that they melt and release synchronously on the same horizontal plane.

[0102] (5) After the dual-line feeding is completed, maintain a weak stirring state (15NL / min) for 8 minutes. After the weak stirring is completed, take a sample to confirm that the N content of the molten steel is within the target range. Further reduce the bottom blowing argon gas to a soft blowing state and maintain it for 5 minutes, then turn off the bottom blowing argon gas.

[0103] (6) The LF-refined steel ladle is transported to the continuous casting turret and cast using a long nozzle argon protection process. The superheat of the molten steel in the tundish is controlled at 25°C, and the steel is continuously cast to obtain duplex stainless steel billets.

[0104] (7) The chemical composition of the obtained duplex stainless steel billet was analyzed. The results showed that the composition of the billet by mass percentage was: C 0.030%, Cr 22.96%, Ni 6.28%, Mo 3.46%, N 0.198%, Si 0.92%, Mn 1.84%, P 0.019%, with the balance being Fe and unavoidable impurities.

[0105] Example 4:

[0106] This embodiment provides a preparation process for a high-purity duplex stainless steel alloy, including the following steps:

[0107] (1) The primary molten steel from the electric arc furnace is added to the AOD furnace, and a bottom-blown O2 / Ar mixed gas is introduced to carry out a decarburization reaction until the C content in the molten steel drops to 0.02%. High-carbon ferrochrome, ferromolybdenum, and pure nickel plates are added for alloying. The amount of pure nickel plates added is determined based on the target Ni content and the amount of Ni brought in by the subsequently fed nickel-magnesium slow-release alloying line. The bottom-blown gas is adjusted to an N2 / Ar mixed gas, and the N content in the molten steel is alloyed to the target upper limit of 0.2% before tapping. The tapping temperature is controlled at 1630℃, and the molten steel is added to the LF refining ladle. No strong deoxidation is carried out during the tapping process.

[0108] (2) After the ladle is positioned in the LF furnace, three-phase electrodes are lowered for heating to stabilize the molten steel temperature at 1600℃. Activated lime and fluorite are added for slag formation, and the binary basicity (CaO / SiO2 mass ratio) of the refining top slag is controlled at 3.8. After the top slag melts, aluminum-manganese-ferroalloy is added for pre-deoxidation, and at the same time, bottom blowing argon gas is turned on in the ladle, with the flow rate set to strong stirring (40 NL / min) and maintained for 6 minutes.

[0109] (3) Under the conditions of maintaining the molten steel temperature at 1600℃ and strong bottom-blown argon stirring (40NL / min), a silicon-barium alloy wire with a barium mass fraction of 12.5% ​​in the powdered core material is fed into the deep part of the molten steel through a wire feeder. The feeding speed is controlled at 2.8m / s, and the amount of barium added is 0.04% of the total mass of the molten steel, based on pure barium element. After the wire feeding is completed, strong stirring (40NL / min) is maintained for another 4 minutes.

[0110] (4) Reduce the bottom-blown argon flow rate to a weak stirring state (12 NL / min). Within 45 seconds of starting the weak stirring, start the dual-wire parallel feeding system and simultaneously feed rare earth cerium-iron alloy wire with a cerium mass fraction of 28.0% in the powdered core material and the nickel-magnesium slow-release alloy wire obtained in Preparation Example 2 into the deep part of the molten steel (400 mm from the bottom of the ladle). The amount of rare earth cerium-iron alloy wire added is 0.015% of the total mass of the molten steel based on pure cerium; the amount of nickel-magnesium slow-release alloy wire added is 0.005% of the total mass of the molten steel based on pure magnesium. Adjust the downward linear velocity of the two alloy wires to 2.0 m / s so that they melt and release synchronously on the same horizontal plane.

[0111] (5) After the dual-line feeding is completed, maintain a weak stirring state (12NL / min) for 6 minutes. After the weak stirring is completed, take a sample to confirm that the N content of the molten steel is within the target range. Further reduce the bottom blowing argon gas to a soft blowing state and maintain it for 5 minutes, then turn off the bottom blowing argon gas.

[0112] (6) The LF-refined steel ladle is transported to the continuous casting turret and cast using a long nozzle argon protection process. The superheat of the molten steel in the tundish is controlled at 22°C, and the steel is continuously cast to obtain duplex stainless steel billets.

[0113] (7) The chemical composition of the obtained duplex stainless steel billet was analyzed. The results showed that the composition of the billet by mass percentage was: C 0.020%, Cr 22.61%, Ni 5.88%, Mo 3.27%, N 0.197%, Si 0.63%, Mn 1.47%, P 0.016%, with the balance being Fe and unavoidable impurities.

[0114] Comparative Example 1:

[0115] Compared with Example 1, the difference is that no silicon-barium alloy wire, rare earth cerium-iron alloy wire, or nickel-magnesium slow-release alloy wire is added. After the pre-deoxidation in step (2), conventional silicon-calcium cored wire is directly fed into the molten steel through a wire feeder. The amount of calcium added is 0.003% of the total mass of the molten steel, based on pure calcium element. Then, long-term weak stirring (bottom-blown argon flow rate 12NL / min, maintained for 20 minutes) is performed to promote gravity floating and impurity removal. All other aspects are the same.

[0116] Comparative Example 2:

[0117] Compared with Example 1, the difference is that the operation of feeding the silicon-barium alloy wire in step (3) and the corresponding stirring time are omitted. After the pre-deoxidation in step (2) is completed, the bottom blowing argon flow rate is directly reduced to a weak stirring state (12NL / min), and then rare earth cerium-iron alloy wire and nickel-magnesium slow-release alloy wire are fed in simultaneously, and the rest are the same.

[0118] Comparative Example 3:

[0119] Compared with Example 1, the difference is that in step (4), only rare earth cerium iron alloy wire is fed in alone, without adding nickel magnesium slow-release alloy wire, that is, the in-situ slow-release micro-air flotation process is missing, while the rest are the same.

[0120] Comparative Example 4:

[0121] Compared with Example 1, the difference is that the core feeding sequence was changed, and the separate feeding and stirring in step (3) was omitted. After the pre-deoxidation was completed, the bottom blowing argon gas was directly reduced to a weak stirring state (12NL / min), and the silicon barium alloy wire, rare earth cerium iron alloy wire and nickel magnesium slow-release alloy wire were fed into the deep part of the molten steel in parallel at the same time. The rest were the same.

[0122] Comparative Example 5:

[0123] Compared with Example 1, the difference is that in step (4), the nickel-magnesium slow-release alloy wire obtained in Preparation Example 2 is replaced with a conventional pure magnesium cored wire (with iron powder as the filler matrix and containing 20.0% pure magnesium powder) with an equal amount of pure magnesium added (0.005% of the total mass of molten steel), and the rest are the same.

[0124] Test Example 1:

[0125] The experimental steps are as follows:

[0126] (1) Weigh about 10.5 mg of the nickel-magnesium alloy powder obtained in Preparation Example 2 and about 10.2 mg of commercially available pure magnesium powder, and place them in a pre-baseline calibrated alumina crucible.

[0127] (2) Place the crucible containing the sample into the heating furnace chamber of the high-temperature synchronous thermal analyzer, seal the furnace body and evacuate it, and then continuously introduce high-purity argon gas, controlling the argon gas flow rate to 50 mL / min to provide an inert protective atmosphere.

[0128] (3) Set the instrument’s heating program to heat from room temperature to 1600℃ at a heating rate of 15℃ / min. Record the mass change and heat flow rate data of the sample in real time during the test. Export the TG curve and DSC curve data after the test.

[0129] The experimental results are shown in Table 1.

[0130] Table 1. TG and DSC test data of nickel-magnesium alloy powder and pure magnesium powder at different temperatures:

[0131]

[0132] According to Table 1 and Figure 1 The data shows that the TG mass retention rate of pure magnesium powder is generally high in the temperature range of 500℃ to 1000℃, but it decreases significantly in the temperature range of 1080℃ to 1150℃, dropping to 12.58% at 1100℃ and further to 0.22% at 1150℃. Simultaneously, its DSC curve shows a large endothermic response in the corresponding temperature range, with a heat flux of -135.62 mW / mg at 1100℃. These results indicate that pure magnesium powder exhibits significant volatilization and weight loss characteristics in the high-temperature range close to its boiling point.

[0133] In contrast, the nickel-magnesium alloy powder obtained in Preparation Example 2 maintained a TG mass retention rate of over 99% within the temperature range of 500℃ to 1150℃, and still reached 93.88% at 1600℃, with a cumulative weight loss rate of 6.12%. Its DSC curve did not exhibit the abrupt, strong endothermic peaks seen in pure magnesium powder throughout the entire test temperature range, and the overall heat flux change was relatively gradual. These results indicate that, under the preparation conditions described in this application, the nickel-magnesium alloy powder exhibits a smoother weight loss behavior and thermal response characteristics compared to pure magnesium powder.

[0134] Based on the above TG and DSC test results, it can be concluded that the release process of magnesium in nickel-based alloy systems is more gradual than that of pure magnesium powder, thus helping to reduce the tendency for concentrated magnesium volatilization under high-temperature conditions. This characteristic can provide experimental evidence for its relatively mild bubble release behavior during steelmaking refining.

[0135] Test Example 2:

[0136] The experimental steps are as follows:

[0137] (1) Block samples with dimensions of 15mm×15mm×20mm were cut from the center of the duplex stainless steel billets formed by continuous casting in Example 1 and Comparative Example 4, respectively. The surface oxide scale was removed by grinding with silicon carbide sandpaper step by step, and the samples were ultrasonically cleaned in anhydrous ethanol and dried with cold air.

[0138] (2) Prepare a mixed solution of anhydrous methanol and acetylacetone with a volume ratio of 10:1, add 1% tetramethylammonium chloride as electrolyte, use the cleaned sample as anode and pure platinum mesh as cathode, place it in an electrolytic cell with a water bath jacket, and control the electrolyte temperature between -5℃ and 0℃.

[0139] (3) The electrolysis process was controlled by a potentiostat. The potential of the anode relative to the saturated calomel electrode was set to -150mV. Electrolysis was continued for 72 hours to completely dissolve the sample matrix metal while retaining the non-metallic inclusions.

[0140] (4) After electrolysis, the anode mud and electrolytic residue are collected into centrifuge tubes and centrifuged at 8000 r / min for 15 minutes. The supernatant is discarded, and then ultrasonic washing and centrifugation are repeated four times with anhydrous methanol.

[0141] (5) Place the purified inclusion powder in a vacuum drying oven and dry it at 60°C for 12 hours. After taking it out, spread it evenly on the glass sample stage.

[0142] (6) The extracted inclusion powder was scanned using an X-ray diffractometer. A Cu-Kα target was used, with a tube voltage of 40 kV and a tube current of 40 mA. The scanning range was set to 2θ = 20° to 80°, the scanning rate was 2° / min, and the step size was 0.02°. The diffraction pattern data was recorded.

[0143] The experimental results are shown in Table 2.

[0144] Table 2. XRD diffraction intensity data of inclusions extracted by electrolysis in Example 1 and Comparative Example 4 at a typical 2θ angle:

[0145]

[0146] According to Table 2 and Figure 2The data from Comparative Example 4 show high diffraction intensities at characteristic 2θ angle positions such as 28.52°, 33.18°, 36.85°, 44.82°, 47.55°, and 59.34°. Specifically, the diffraction intensities at 36.85°, 44.82°, and 59.34° correspond to the expected crystalline phase of magnesium aluminum spinel (MgAl2O4) at 2145 a.u., 1632 a.u., and 1488 a.u., respectively; at 28.52° and 47.55°, the diffraction intensities at 28.52° and 47.55° correspond to the expected crystalline phase of rare earth oxide (CeO2) at 1287 a.u. and 1056 a.u., respectively; and at 33.18°, the diffraction intensity at 33.18° corresponds to the expected crystalline phase of rare earth oxide (Ce2O3) at 893 a.u. These results indicate that, under the test conditions, Comparative Example 4 exhibits relatively clear characteristics of the relevant crystalline phases extracted from the inclusions.

[0147] In contrast, the diffraction intensities of Example 1 at the same characteristic 2θ angle positions were 215 a.u., 194 a.u., 305 a.u., 278 a.u., 212 a.u., and 201 a.u., respectively, which were significantly lower than those of Comparative Example 4. At the same time, the intensity differences between the background positions such as 31.05°, 42.11°, 52.40°, and 65.22° and the characteristic peak intensity were relatively small, indicating that the sharpness of the diffraction peaks of the inclusions extracted in Example 1 was relatively weakened, and the crystalline phase characteristics were not as prominent as those in Comparative Example 4.

[0148] Based on the results shown in Table 2, it can be concluded that under the process conditions described in this application, the characteristic peaks of the highly crystalline magnesium aluminum spinel phase and rare earth oxide phase in the inclusions obtained in Example 1 were significantly weakened. This difference suggests that the stepwise sequential addition method may be beneficial in reducing the tendency of the relevant inclusions to develop into regular crystal phase morphology, thereby providing favorable conditions for the subsequent flotation and removal of inclusions.

[0149] Test Example 3:

[0150] The experimental steps are as follows:

[0151] (1) At the industrial LF furnace refining and continuous casting rotary table operation site, the metallurgical parameters of each furnace corresponding to Examples 1 to 4 and Comparative Examples 1 to 5 were tracked and recorded. The time node from the completion of the pre-deoxidation process to the start of the long nozzle protection pouring was taken as the data collection interval of the refining process.

[0152] (2) A continuous temperature measuring thermocouple is periodically inserted into the ladle to monitor the initial temperature of the molten steel when it is tapped from AOD and enters the LF furnace. Then, the ladle is received on the continuous casting rotary table and the final pouring temperature is measured. The difference between the two during this period is calculated to obtain the total temperature drop data from tapping from AOD to pouring.

[0153] (3) The operation time of the weak stirring stage is timed, and the actual number of minutes consumed from the end of the last wire feeding to the end of the impurity removal and the achievement of the target total oxygen content standard is recorded. This time is taken as the weak stirring and impurity removal time after wire feeding. Observe and record the metallurgical reaction phenomena on the surface of the molten steel during the wire feeding operation. Special records are made for heats in which abnormal conditions such as splashing occur.

[0154] (4) Using quartz sampling tubes, deep molten steel was vacuum-absorbed at the tapping point of the AOD furnace and inside the continuous casting tundish. The obtained samples were quenched in water, cooled, and polished. The mass percentage of nitrogen was determined using a direct-reading spectrometer and a high-precision oxygen-nitrogen analyzer. The nitrogen retention rate of this heat was calculated by comparing the nitrogen content of the molten steel sample from the continuous casting tundish with the nitrogen content of the AOD tapping sample.

[0155] The experimental results are shown in Table 3. For each experimental group, the duration of weak stirring after the wire feeding process ended was defined as the time until the target total oxygen content standard was reached. The time listed in Table 3 is the actual time consumed for the corresponding furnace. For furnaces that experienced abnormal interruptions, the time listed in Table 3 is the actual duration from the end of the last wire feeding to the forced termination of the process.

[0156] Table 3. Results of refining parameters and nitrogen retention index measurements for each experimental group:

[0157]

[0158] According to the data in Table 3, no obvious abnormalities such as splashing occurred during the metallurgical operation in Examples 1-4. The weak stirring and impurity removal time after wire feeding was 5.3 min to 8.1 min. The total temperature drop from AOD tapping to casting was 19.8℃ to 24.2℃. The nitrogen content in the continuous casting tundish steel was 0.177wt% to 0.198wt%, corresponding to a nitrogen retention rate of 97.89% to 99.00%. The above results indicate that under the process conditions of this application, the example group maintained a high nitrogen retention level while the duration of the impurity removal process and the temperature drop were both within a low range.

[0159] In comparison, the weak stirring and impurity removal time after wire feeding treatment in Comparative Examples 1–4 was 14.5 min–20.5 min, the total temperature drop from AOD tapping to casting was 39.8℃–52.3℃, and the nitrogen retention rate was 84.04%–90.26%, all of which were inferior to the Example Group. Among them, Comparative Examples 1 and 3, lacking corresponding modification or air flotation conditions, required significantly longer weak stirring times to reach the target total oxygen content standard; although the process flow of Comparative Examples 2 and 4 was adjusted, their impurity removal time, temperature drop, and nitrogen retention rate still did not reach the level of the Example Group.

[0160] In Comparative Example 5, the duration of weak stirring after wire feeding was 4.5 minutes. However, this heat experienced violent boiling and splashing during refining, which forced the weak stirring to be interrupted. Therefore, its time data should not be equated with heats that normally complete impurity removal. The total temperature drop from AOD tapping to casting start in this heat was 61.4℃, the nitrogen content in the continuous casting tundish was 0.142 wt%, and the nitrogen retention rate was 75.13%, indicating poor process stability under these conditions.

[0161] Based on the results shown in Table 3, it can be concluded that the process conditions adopted in the embodiments of this application are beneficial to shorten the impurity removal operation time, reduce the temperature drop, and improve the nitrogen retention level; however, when the key modification steps are omitted, the air flotation carrier is omitted, or conventional pure magnesium cored wire is used, the process stability and nitrogen retention effect are reduced.

[0162] Test Example 4:

[0163] The experimental steps are as follows:

[0164] (1) Metal blocks were cut from the cross-section of the continuous casting duplex stainless steel billets obtained in Examples 1-4 and Comparative Examples 1-5 at 1 / 4 thickness and the center, respectively, and machined into cylindrical analytical samples with a diameter of 4 mm and a weight of about 1.0 g using a lathe.

[0165] (2) Use a special fine file to grind the surface of the cylindrical sample to completely remove the surface oxide layer and contaminants that may be introduced during the processing. Then, put it into a beaker containing anhydrous acetone for ultrasonic cleaning for 5 minutes. After taking it out, blow it with cold air until it is completely dry and put it in a desiccator for later use.

[0166] (3) The total oxygen content (TO) in the sample was determined using a pulse-heated inert gas melting-infrared absorption spectrometer. A quartz crucible was placed between the upper and lower electrodes for high-temperature degassing and air-firing. After cooling, the pretreated analytical sample was added, and a high-current pulse heating was applied under a helium carrier gas flow to melt the sample. The carbon monoxide and carbon dioxide generated by the reaction of oxygen in the sample with the graphite crucible entered the infrared detection cell with the carrier gas. The total oxygen mass fraction was calculated and output using integration software and converted to a value at the ppm level.

[0167] (4) The total sulfur content (TS) in the sample was determined using a high-frequency combustion infrared absorption spectrometer. Another set of pretreated samples was placed in a ceramic crucible, and about 1.5g of high-purity tungsten granules and 0.3g of pure tin granules were added as flux. The crucible was then placed in a high-frequency induction furnace and high-purity oxygen was introduced for combustion. After the sulfur in the sample was completely converted into sulfur dioxide gas, it was measured in an infrared absorption cell. The software automatically converted the total sulfur content to ppm level values.

[0168] (5) Five parallel samples were measured for each experimental group. After removing outliers, the arithmetic mean was calculated as the final macroscopic purity index of the billet in that furnace.

[0169] The experimental results are shown in Table 4.

[0170] Table 4. Results of determination of total oxygen (TO) and total sulfur (TS) content in the final cast slabs of the Examples and Comparative Examples:

[0171]

[0172] According to the data in Table 4, the total oxygen content of the continuously cast duplex stainless steel billets obtained in Examples 1-4 was 6.5 ppm to 7.8 ppm, and the total sulfur content was 3.5 ppm to 4.6 ppm, with a data fluctuation deviation rate of ±3.8% to ±5.5%. The above results indicate that under the process conditions of this application, the total oxygen and total sulfur contents in the billets of the example groups remained at low levels, and the data fluctuations within the groups were small, indicating good process repeatability.

[0173] In contrast, the total oxygen content of Comparative Examples 1–4 ranged from 12.5 ppm to 15.4 ppm, and the total sulfur content ranged from 8.7 ppm to 12.1 ppm, with a data fluctuation deviation of ±7.4% to ±11.2%, all of which were higher than those of the Example Group. This result indicates that in the absence of corresponding sequential modification steps, barium-based liquid phase construction steps, or slow-release flotation conditions, the removal efficiency of deoxidation and desulfurization products in molten steel decreases, leading to an increase in the total oxygen and total sulfur content in the final cast billet.

[0174] Comparative Example 5 had a total oxygen content of 21.6 ppm and a total sulfur content of 15.4 ppm, with a data fluctuation deviation rate of ±18.6%, the highest among all experimental groups. Considering the process anomalies observed in this group during the aforementioned tests, it can be concluded that under conventional pure magnesium cored wire conditions, the process stability is relatively poor, and the final billet purity control effect is not as good as that of the Example Group.

[0175] Based on the results shown in Table 4, it can be concluded that the process conditions adopted in the embodiments of this application are beneficial to reducing the total oxygen and total sulfur content in the final cast billet, and improving process stability and result consistency.

[0176] Test Example 5:

[0177] The experimental steps are as follows:

[0178] (1) A block sample with dimensions of 10mm×10mm×5mm was cut from the center of the duplex stainless steel billet continuously cast in Example 1 and Comparative Example 1. After welding copper wires to the back of the sample, it was cold-mounted with epoxy resin, leaving an area of ​​1cm². 2 The working test surface.

[0179] (2) The working surface of the sample was polished step by step to 2000 grit using wet sandpaper, and then mechanically polished to a mirror state using diamond polishing liquid with a particle size of 1.0 μm. After being cleaned and degreased by ultrasonic cleaning with anhydrous ethanol, it was dried with cold air and placed in a desiccator for later use.

[0180] (3) The constant potential polarization temperature rise test was carried out in accordance with the ASTM G150 standard. A sodium chloride (NaCl) aqueous solution with a concentration of 3M was prepared as the electrolyte and transferred into an electrolytic cell with a constant temperature water bath jacket. Before the test, high-purity nitrogen was continuously introduced into the solution to remove oxygen and degas for 1 hour, and a nitrogen protective micro-positive pressure was maintained above the liquid surface throughout the test.

[0181] (4) Construct a standard three-electrode test system, with the embedded duplex stainless steel sample as the working electrode, the saturated calomel electrode (SCE) as the reference electrode, and the large-area platinum mesh as the auxiliary electrode.

[0182] (5) A constant anodic potential of +700mV (vs. SCE) was applied to the working electrode using an electrochemical workstation. After stabilizing the polarization for 10 minutes, a programmable temperature-controlled water bath was started to linearly heat the electrolyte at a set rate of 1℃ / min. The response current density on the surface of the working electrode was recorded synchronously throughout the test. When the current density of the test system continued to rise and exceeded 100μA / cm 2 When the threshold is reached, the corresponding temperature is recorded as the critical pitting temperature (CPT), and then the test program is terminated.

[0183] The experimental results are shown in Table 5.

[0184] Table 5. Current density variation data of Example 1 and Comparative Example 1 during constant potential polarization heating process:

[0185]

[0186] According to the data in Table 5, the current density of Comparative Example 1 in the range of 30.2℃ to 47.1℃ is 2.14 μA / cm². 2 It gradually increased to 18.24 μA / cm 2 When the temperature rises to 48.4℃, its current density reaches 105.32 μA / cm². 2 Exceeding 100 μA / cm 2 The threshold, corresponding to the critical pitting temperature (CPT), is 48.4 °C. Subsequently, the current density further increases to 285.47 μA / cm² at 49.5 °C. 2 And it reaches the upper limit of the instrument's measuring range at 55℃.

[0187] In contrast, the current density in Example 1 remained at a low level throughout the temperature range of 30.2°C to 67.3°C, at 1.85 μA / cm².2 It gradually increased to 19.45 μA / cm 2 When the temperature rises to 68.2℃, its current density reaches 108.62 μA / cm². 2 Exceeding 100 μA / cm 2 The threshold, corresponding to the critical pitting temperature (CPT), is 68.2 °C. Subsequently, the current density further increases to 315.84 μA / cm² at 69.5 °C. 2 It reaches the upper limit of the instrument's measuring range at 71℃.

[0188] The above results indicate that, under the test conditions, the critical pitting temperature of Example 1 is higher than that of Comparative Example 1, suggesting that its resistance to localized corrosion in chlorine-containing media is relatively better. Combined with the aforementioned inclusion control results, it can be concluded that the refining process used in the embodiments of this application is beneficial for improving the pitting corrosion resistance of the material.

[0189] Although embodiments of the invention have been shown and described, it will be understood by those skilled in the art that various changes, modifications, substitutions and alterations can be made to these embodiments without departing from the principles and spirit of the invention, the scope of which is defined by the appended claims and their equivalents.

Claims

1. A preparation process for a high-purity duplex stainless steel alloy, characterized in that, Includes the following steps: S1. The primary steel molten steel is poured into the AOD furnace, bottom-blown mixed gas is introduced to carry out decarburization reaction, alloy is added to carry out alloying, the tapping temperature is controlled and the steel molten steel is poured into the LF refining ladle, and strong deoxidation is not carried out during the tapping process. S2. After the ladle is placed in the LF furnace, the electrodes are lowered for heating, slag-forming material is added for slag formation, aluminum-manganese-iron alloy is added for pre-deoxidation, and at the same time, bottom blowing argon gas is turned on in the ladle, so that the bottom blowing argon gas flow rate is 35NL / min~45NL / min and a strong stirring state is formed, and it is maintained for 5~8 minutes. S3. While maintaining strong stirring, feed the silicon-barium alloy wire into the deep part of the molten steel. After feeding the wire, continue to maintain strong stirring for 3 to 5 minutes. S4. Reduce the bottom-blown argon flow rate to 10NL / min~15NL / min to create a weak stirring state. Within 30~60 seconds after the weak stirring state is established, start the dual-line parallel feeding system to simultaneously feed rare earth cerium-iron alloy wire and nickel-magnesium slow-release alloy wire into the depth of the molten steel, so that the two alloy wires melt and release synchronously on the same horizontal plane. Adjust the downward linear velocity of the two alloy wires to 1.5m / s~2.5m / s. Among them, based on pure cerium, the amount of rare earth cerium-iron alloy wire added is 0.01%~0.02% of the total mass of the molten steel; based on pure magnesium, the amount of nickel-magnesium slow-release alloy wire added is 0.003%~0.008% of the total mass of the molten steel. S5. After the double-line feeding is completed, maintain the air flotation and impurity removal under weak stirring for 5 to 8 minutes. Then, further reduce the bottom blowing argon gas to a soft blowing state and maintain it for 5 minutes before turning off the bottom blowing argon gas. S6. The refined steel ladle is transported to the continuous casting turret and continuously cast using a long nozzle argon-protected casting process to obtain a duplex stainless steel billet. The duplex stainless steel alloy obtained comprises the following components by mass percentage: C: 0.015%–0.030%, Cr: 22.0%–23.0%, Ni: 4.5%–6.5%, Mo: 3.0%–3.5%, N: 0.14%–0.20%, Si: 0.3%–1.0%, Mn: 0.8%–2.0%, P≤0.02%, with the balance being Fe and unavoidable impurities.

2. The preparation process according to claim 1, characterized in that, The silicon-barium alloy wire is made of a powdered core material wrapped in an outer layer of low-carbon steel strip. The powdered core material contains 10.0% to 15.0% barium and 40.0% to 50.0% silicon by mass. The rare earth cerium-iron alloy wire is made of a powdered core material wrapped in an outer layer of low carbon steel strip, and the powdered core material contains 25.0% to 30.0% cerium by mass. The nickel-magnesium slow-release alloy wire is made of nickel-magnesium alloy powder coated with an outer layer of cold-rolled low-carbon steel strip. The particle size of the nickel-magnesium alloy powder is 75μm to 150μm, and the nickel-magnesium alloy powder is composed of the following components by mass percentage: magnesium 4.0% to 8.0%, with the balance being nickel and unavoidable impurities.

3. The preparation process according to claim 1, characterized in that, In step S1, a decarburization reaction is carried out until the C content in the molten steel drops to 0.015%–0.030%; the bottom blowing gas is adjusted to a N2 / Ar mixed gas to alloy the N content to 0.18%–0.20%; the tapping temperature is controlled at 1620℃–1640℃. In step S2, heating is controlled to stabilize the temperature of the molten steel at 1590℃~1610℃; the mass ratio of the binary basicity CaO / SiO2 in the refining top slag is controlled to be 3.5~4.

0.

4. The preparation process according to claim 1, characterized in that, In step S3, the feed rate of the silicon-barium alloy wire is 2.5 m / s to 3.0 m / s, and the amount of barium added is 0.03% to 0.05% of the total mass of the molten steel, based on pure barium element.

5. The preparation process according to claim 1, characterized in that, In step S6, the superheat of the molten steel in the continuous casting tundish is controlled between 20°C and 25°C.

6. The preparation process according to claim 1, characterized in that, The nickel-magnesium slow-release alloy wire used in step S4 is prepared in advance by the following process: (1) Calculate the amount of raw materials according to the target alloy mass fraction of 4.0% to 8.0% magnesium and 92.0% to 96.0% nickel, and add an extra pure magnesium block as compensation for burn-off; (2) Place the pure nickel block in the crucible of the vacuum induction melting furnace, evacuate the vacuum, and heat up to completely melt the pure nickel block to obtain pure nickel liquid. (3) High-purity argon gas is introduced into the melting furnace, and the pure magnesium block is quickly pressed into the bottom of the pure nickel liquid using a perforated refractory bell jar, while maintaining induction electromagnetic stirring to fully alloy it. (4) The obtained alloy liquid is cast into a water-cooled copper mold and cooled to obtain a nickel-magnesium alloy ingot; after coarse crushing, it is placed in a ball mill and ground under argon protection, and nickel-magnesium alloy powder with a particle size of 75μm~150μm is collected by sieving. (5) The cold-rolled low-carbon steel strip is bent into a U-shaped groove by the unit, the nickel-magnesium alloy powder is added, closed and compacted, and then drawn into a cored wire.

7. The preparation process according to claim 1, characterized in that, The obtained duplex stainless steel alloy has a total oxygen content of 6.5ppm to 7.8ppm and a total sulfur content of 3.5ppm to 4.6ppm.

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