High-strength alloy steel for shield cutter and heat treatment method thereof

Abstract

The application relates to the field of alloy steel materials and metal heat treatment technology, and discloses a high-strength alloy steel for a shield cutter and a heat treatment method of the high-strength alloy steel. The alloy steel is composed of 0.55-0.65 wt% of C, 1.2-1.6 wt% of Mn, 2.5-3.5 wt% of Mo, W, V, Ti, Zr and B. When the steel liquid with the above chemical components is prepared, the steel liquid is vacuum degassed and refined, then is fed into low-silicon aluminum zirconium cored wire for deoxidization to reduce free oxygen activity, and then titanium-iron alloy and boron-iron cored wire are sequentially added; after forming and processing, the component is sequentially subjected to double-stage solid solution treatment, salt bath isothermal grading quenching, in-situ deep cooling treatment and double-stage targeted tempering. The application protects the solid solution precipitation state of micro-alloy elements by time sequence feeding, combines with multi-stage heat treatment to control the section temperature difference and carbide precipitation, improves the hardenability and the matching of strength and toughness of the material, and is suitable for the manufacturing of the shield cutter.

Description

Technical Field

[0001] This invention relates to the field of alloy steel materials and metal heat treatment technology, and in particular to a high-strength alloy steel for tunnel boring machine tools and its heat treatment method. Background Technology

[0002] Tunnel boring machines (TBMs) are widely used in the construction of subway tunnels, cross-river tunnels, and underground utility tunnels. As the core component of the TBM that directly contacts the working face and breaks up rock and soil, the cutting tools typically operate in harsh environments. When cutting through geological layers containing high-hardness minerals such as quartz and feldspar, the outer surface of the cutting tools is subjected to intense abrasive erosion. Furthermore, when encountering uneven rock formations or geological fault zones, the cutting edge must withstand extremely high alternating impact loads. These working conditions require the cutting tool steel to possess not only high macroscopic hardness to resist wear but also sufficient toughness to prevent chipping or premature fracture.

[0003] Currently, the materials commonly used in manufacturing tunnel boring machine (TBM) cutters are mainly medium- and high-carbon alloy structural steels. To improve the overall hardenability of thick-section cutter components, microalloying elements are often introduced into the molten steel. Boron is often chosen as the preferred additive element to improve hardenability because it can effectively delay the nucleation of proeutectoid ferrite along grain boundaries at extremely low concentrations. However, during conventional smelting, free boron is chemically active and readily reacts with residual oxygen and nitrogen in the molten steel to form boron oxide or boron nitride. This reaction not only consumes the effective boron content originally used for solid solution strengthening, weakening its actual hardenability, but also inevitably increases the number of non-metallic inclusions inside the steel, thus becoming stress concentration points when the matrix is ​​under stress.

[0004] In the post-forming phase transformation heat treatment process, conventional single rapid quenching often creates a significant temperature difference between the surface and the core of the workpiece when processing tools with large cross-sectional dimensions. The uneven volume shrinkage and asynchronous martensitic phase transformation caused by this temperature difference lead to the accumulation of substantial residual thermal stress and microstructure transformation stress, increasing the risk of early tool cracking. While some existing processes attempt to release stress by adjusting the tempering temperature, they often struggle to simultaneously address the fine dispersion and precipitation of microscopic carbides. This creates a trade-off in the material's mechanical properties: prioritizing hardness increases brittleness and reduces impact resistance; conversely, prioritizing toughness decreases resistance to abrasive wear. Therefore, finding a more reasonable balance between strength and toughness through coordinated control from the initial refining and feeding sequence to the subsequent thermodynamic phase transformation process is a crucial technical requirement for improving tool life. Summary of the Invention

[0005] The technical problem to be solved by this invention is that shield tunneling cutters face abrasive erosion and impact loads when breaking rocks in complex geological conditions, while conventional alloy steel materials often have insufficient hardenability or difficulty in achieving both strength and toughness.

[0006] To address the above problems, the present invention provides the following technical solution: In a first aspect, the present invention provides a high-strength alloy steel for tunnel boring machine cutting tools, which is composed of the following chemical components by mass percentage: C: 0.55~0.65wt%; Mn: 1.2~1.6wt%; Mo: 2.5~3.5wt%; W: 1.5~2.0wt% V: 1.0~1.5wt%; Ti: 0.05~0.1wt%; Zr: 0.015~0.025wt%; B: 0.003~0.006wt%; Si: 0.03~0.08wt% Ni: 0.02~0.08wt% Cr: 0.02~0.08wt%; S: ≤0.004wt% P: ≤0.008wt% The balance consists of Fe and unavoidable impurities; The alloy steel is made by degassing and refining molten steel containing the above-mentioned chemical components in a vacuum, then feeding low-silicon aluminum zirconium iron cored wire with low-silicon aluminum zirconium iron alloy powder as the core powder to deoxidize and reduce the free oxygen activity of the molten steel. Subsequently, titanium iron alloy and boron iron cored wire are added in sequence. After casting and forging, the steel is then subjected to two-stage solution treatment, salt bath isothermal graded quenching, in-situ deep cryogenic treatment, and two-stage targeted tempering.

[0007] By adopting the above technical solution and combining it with a specific chemical composition ratio to implement sequential microalloying and multi-stage phase transformation heat treatment, the following effects are achieved: Adding a small amount of boron can generally improve the hardenability of alloy steel; however, free boron tends to react with oxygen and nitrogen in the molten steel to form inclusions, affecting the actual microalloying effect. Considering the final state of the dissolved boron, the scheme adopts a sequential refining feeding design.

[0008] Specifically, after vacuum degassing, low-silicon aluminum zirconium iron cored wire is fed into the molten steel. Because zirconium has a strong chemical affinity for oxygen, after the alloy powder is fed into the deep part of the molten steel using the cored wire for melting, zirconium will undergo a deoxidation reaction with free oxygen. [Zr]+2[O]→ZrO2; Most of the generated zirconium oxide particles float to the slag layer and are discharged, while a small number remain in the molten steel to serve as heterogeneous nucleation sites during subsequent solidification. The deoxidation process reduces the free oxygen activity in the molten steel, thereby inhibiting the tendency of subsequently added boron to be oxidized to some extent. In addition, using a low-silicon, low-alumina core powder formulation also helps control the increase in brittle inclusions.

[0009] After the free oxygen activity is reduced, titanium-iron alloy is then added. The added titanium mainly serves to bind residual free nitrogen in the molten steel, resulting in the following reaction: [Ti]+[N]→TiN; As titanium nitride particles precipitate within the matrix, free nitrogen is further consumed, which reduces the likelihood of boron nitride formation during subsequent boron addition.

[0010] The preceding deoxidation and nitrogen fixation treatments improved the cleanliness of the molten steel before the ferroboron-boron cored wire was fed in. At this stage, the added boron is largely retained in the molten steel in a solid solution state. During the subsequent solidification and phase transformation stages, the solid-solidified boron tends to accumulate at the austenite grain boundaries and lower the grain boundary energy. This often delays the nucleation of proeutectoid ferrite along the grain boundaries, thus positively impacting the hardenability of the alloy steel.

[0011] During the forming and heat treatment stages, the alloy material undergoes microstructural transformation through multi-stage thermo-mechanical coupling. Two-stage solution treatment facilitates the dissolution of carbide-forming elements into the austenite matrix while simultaneously constraining the austenite grain size. Subsequent isothermal graded salt bath quenching reduces the temperature difference across the workpiece cross-section, promoting the transformation of austenite into martensite and lower bainite to alleviate transformation stress. The subsequent in-situ cryogenic treatment further promotes the phase transformation of residual austenite. Finally, two-stage targeted tempering promotes the precipitation of nanoscale carbides, resulting in secondary hardening and further adjusting the balance between hardness and toughness of the alloy steel.

[0012] Preferably, the alloy steel is composed of the following chemical components by mass percentage: C: 0.60 wt%, Mn: 1.4 wt%, Mo: 3.0 wt%, W: 1.75 wt%, V: 1.25 wt%, Ti: 0.075 wt%, Zr: 0.02 wt%, B: 0.0045 wt%, Si: 0.05 wt%, Ni: 0.05 wt%, Cr: 0.05 wt%, S: 0.003 wt%, P: 0.005 wt%, with the balance being Fe and unavoidable impurities.

[0013] By adopting the above technical solution, when carbon and carbide-forming elements are in the preferred ratio, the material tends to precipitate composite carbides with a suitable volume fraction during the tempering stage, while the matrix maintains a certain amount of carbon solid solution. This microstructural regulation helps to coordinate the material's resistance to abrasive wear and impact fracture, which is beneficial to improving the actual service life of tool components.

[0014] Secondly, the present invention provides a method for preparing high-strength alloy steel for tunnel boring machine cutters, comprising the following steps: (1) Batching and primary refining: Based on the target chemical composition, the raw materials are calculated and weighed, the bottom material is added to the electric arc furnace for melting, oxygen blowing for decarburization, slag formation for dephosphorization, and then the steel is tapped to the ladle; (2) Refining and sequential feeding: The ladle is transferred to the LF furnace for desulfurization and basic alloying, and then transferred to the VD vacuum degassing furnace for deep degassing; before breaking the vacuum, after confirming that the free oxygen activity of the molten steel meets the standard, low silicon aluminum zirconium iron cored wire, titanium iron alloy, and boron iron cored wire are fed into the molten steel in sequence, and alloy steel ingots are obtained by die casting after breaking the vacuum. (3) Forging and annealing: After heating and holding the alloy steel ingot, it is forged in multiple directions, then annealed and slowly cooled to form a tool blank; (4) Two-stage solution treatment: The tool blank is heated and kept at a certain temperature in a vacuum heat treatment furnace for the first stage of solution treatment; then the heating is stopped and high-purity nitrogen is introduced to cool down and keep it at a certain temperature for the second stage of solution treatment. (5) Salt bath isothermal graded quenching: After completing the two-stage solution treatment, the knife blank is quickly transferred and immersed in a mixed molten nitrate bath for isothermal holding. After taking it out, it is air-cooled to room temperature. (6) In-situ cryogenic treatment: The blank is transferred into a cryogenic chamber, cooled and kept warm with liquid nitrogen, and then slowly and naturally warmed to room temperature; (7) Two-stage targeted tempering: The shield cutter is made of high-strength alloy steel by going through two heating, holding and air cooling steps at different temperatures.

[0015] By adopting the above technical solutions, a process control chain for high-alloy systems has been formed from the smelting front end to the heat treatment end, which is described in detail below: In the smelting and forming processes, sulfur, phosphorus, and gaseous impurities in the molten steel are first reduced through primary smelting and vacuum refining. The sequential charging operation utilizes the aforementioned stepwise deoxidation and nitrogen fixation mechanism, effectively protecting microalloying elements such as boron by controlling the timing and method of addition. In the subsequent forming stage, multi-directional forging helps break up the coarse dendrites in the as-cast microstructure, mitigating compositional segregation to some extent and providing a relatively uniform matrix structure for subsequent heat treatment.

[0016] In the phase transformation heat treatment process, a multi-stage control method was introduced. The solution treatment was carried out in two temperature steps. The first stage of high-temperature heating caused most of the carbide-forming elements to dissolve into the austenite. Subsequently, instead of quenching directly, high-purity nitrogen was introduced to slowly lower the furnace temperature to the second stage temperature for a short-term holding. This stepped cooling process can often fine-tune the carbon concentration distribution in the austenite before quenching, while reducing the temperature difference between the workpiece surface and the core, thereby reducing the risk of excessive thermal stress during subsequent rapid cooling.

[0017] The subsequent quenching step employed isothermal staged quenching in a salt bath. Utilizing the relatively mild cooling and heat transfer characteristics of the nitrate salt bath, the temperature difference across the cross-section was further controlled, promoting the composite transformation of supercooled austenite into martensite and lower bainite. This helps improve the material's toughness while maintaining high hardness. The subsequent cryogenic treatment used liquid nitrogen to further induce phase transformation in the retained austenite. Two stages of targeted tempering at different temperatures gradually released the quenching stress and promoted the dispersed precipitation of alloy carbides from the martensitic matrix, thus establishing the final mechanical properties of the material.

[0018] Preferably, in step (1), the base material consists of 70-80 wt% direct reduced iron and 20-30 wt% carbon structural steel scrap. During the initial refining process, graphite carbonizer is added to adjust the carbon content according to the target carbon content. In step (2), the temperature of the molten steel is adjusted to 1620-1650℃ in the LF furnace, and the vacuum degree of the VD vacuum degassing furnace is maintained at 40-60Pa for 15-20 minutes for deep degassing. 3-5 minutes before breaking the vacuum, the free oxygen activity of the molten steel is confirmed to be 10-15ppm. Low silicon aluminum zirconium iron cored wire is fed in at a speed of 2.0-3.0m / s and weakly stirred by argon blowing at the bottom and left to stand for 1-2 minutes. Titanium iron alloy is added and weakly stirred by argon blowing for 1 minute. Finally, boron iron cored wire is fed in and argon blowing for 1 minute and then molded at 1530-1550℃. The low-silicon aluminum zirconium iron cored wire described in step (2) is made by mixing 40-45 wt% zirconium powder and 55-60 wt% iron powder, heating the mixture to 1550-1600℃ under argon protection in a yttrium oxide crucible of a medium-frequency induction furnace until fully melted; after natural cooling, the mixture is crushed and sieved to collect 60-80 mesh alloy powder as core powder, which is then continuously wrapped and formed using low-carbon steel strip. The silicon mass fraction in the core powder is <1.0 wt%, and the aluminum mass fraction is <0.5 wt%. In step (3), the alloy steel ingot is heated to 1150-1180℃ at a rate of 15℃ / min and held for 2-3 hours. It is then forged in multiple directions at 900-1050℃ with a total forging ratio of 6-8. After forging, it is air-cooled. It is then heated to 820-840℃ for annealing and held for 4-5 hours. Finally, it is cooled in the furnace at a rate of 15-20℃ / h to 500-520℃ and then air-cooled. In step (4), the vacuum is evacuated to 0.02-0.08 Pa. In the first stage, the temperature is heated to 1060-1080°C at a rate of 10-15°C / min and held for 1.5-2 hours. In the second stage, high-purity nitrogen is introduced and the temperature is cooled to 920-950°C at a rate of 3-5°C / min and held for 30-40 minutes. In step (5), the blade blank is immersed in a molten nitrate bath containing 50wt% potassium nitrate and 50wt% sodium nitrite at a temperature of 280-290°C for 1.5-2 hours within 15 seconds; In step (6), the sample is transferred to a cryogenic chamber within 1 hour and kept at -120 to -130°C for 4 to 5 hours at a rate of 2.0 to 3.0°C / min. In step (7), the first stage involves heating to 520-530℃ and holding for 2-3 hours before air cooling, while the second stage involves heating to 550-560℃ and holding for 2-3 hours before air cooling.

[0019] By adopting the above technical solution and introducing a higher proportion of direct reduced iron during the batching process, the amount of residual harmful impurities introduced can be controlled from the smelting source, thereby improving the purity of the raw material. Limiting the specific free oxygen activity index and wire feeding speed during the refining process helps the cored wire overcome the buoyancy of the molten steel and high-temperature consumption, allowing it to penetrate directly to the lower part of the molten steel for release, thus maintaining the stability of the deoxidation and microalloying reactions.

[0020] Controlling the preparation parameters of low-silicon aluminum zirconium iron cored wire, especially by using yttrium oxide crucible for melting and limiting the silicon and aluminum content, greatly reduces the possibility of introducing silicate and alumina inclusions that easily form brittle phases into the molten steel.

[0021] At the level of heat treatment parameters, limiting the heating, holding time, and cooling rate to the aforementioned range comprehensively considers the dissolution kinetics and precipitation thermodynamics of alloy carbides. A suitable forging ratio combined with annealing prepares the microstructure conditions for subsequent phase transformation; precise control of the two-stage solution treatment and graded quenching temperatures balances the relationship between grain growth tendency and microstructure transformation depth, providing good process support for ultimately improving the service life of tunnel boring machine cutter components in the field.

[0022] In summary, the present invention has at least one of the following beneficial technical effects: 1. This invention improves the hardenability and internal microstructure purity of alloy steel by combining specific chemical composition ratios with a sequential microalloying process. After a vacuum deep degassing process, custom-designed low-silicon aluminum zirconium iron cored wire, titanium iron alloy, and boron iron cored wire are introduced sequentially according to a specific time sequence. The deep deoxidation effect of zirconium and the nitrogen-fixing effect of titanium create a low-oxygen, low-nitrogen environment in the molten steel for the subsequent addition of boron. This step-by-step feeding mechanism effectively reduces the oxidation loss of boron, promoting its stable existence in a solid solution state and enrichment at austenite grain boundaries, delaying the nucleation of proeutectoid ferrite, and thus improving the overall hardenability of the alloy material. Simultaneously, the use of a low-silicon, low-aluminum core powder formula limits the formation of brittle inclusions such as silicates to a certain extent, which is beneficial for improving the macroscopic impact resistance of the matrix.

[0023] 2. This invention further adjusts the strength and toughness matching of tool components by combining the smelted and shaped material with a multi-stage phase transformation heat treatment process. The two-stage solution treatment utilizes stepped cooling to fine-tune the carbon concentration distribution within the austenite body and reduces the cross-sectional temperature difference of the workpiece before quenching, lowering the risk of excessive thermal stress during the subsequent rapid cooling stage. Subsequent isothermal graded quenching in a nitrate bath combined with in-situ cryogenic treatment promotes the transformation of the matrix structure into a composite structure of martensite and lower bainite. Based on this, a two-stage targeted tempering process promotes the dispersion and precipitation of nanoscale composite carbides, resulting in a secondary hardening effect. The alloy material prepared using this process maintains high hardness to resist abrasive wear while moderately alleviating internal phase transformation stress, often exhibiting good structural stability and service life when dealing with the high-load impact conditions during shield tunneling rock breaking. Attached Figure Description

[0024] Figure 1 This is a schematic diagram of the representative thermal expansion response and local X-ray diffraction characteristics in Test Example 1 of the present invention; wherein, (a) is a schematic curve of the thermal expansion response of Example 1, Comparative Example 4 and Comparative Example 5 in the isothermal stage; (b) is a schematic diagram of the local X-ray diffraction characteristics of the corresponding sample in the range of 2θ=42°~46° after heat treatment. Figure 2 This is a schematic diagram of the local X-ray diffraction characteristics of the representative extract phase in Test Example 2 of the present invention; wherein, (a) is a schematic diagram of the local characteristic peaks of Example 1, Comparative Example 3 and Comparative Example 1 in the range of 2θ=33°~44°; (b) is a schematic diagram of the local characteristic peaks of Example 1, Comparative Example 3 and Comparative Example 1 in the range of 2θ=50°~72°. Figure 3 This is a comparison curve of the coefficient of friction as a function of sliding distance in Example 1, Comparative Example 1 and Comparative Example 4 of Test Example 4 of the present invention under dry sliding friction conditions; Figure 4This is a comparison curve of the dynamic stiffness of the system under high-frequency impact load as a function of the number of impact cycles in Example 1, Comparative Example 1 and Comparative Example 5 of the present invention. Detailed Implementation

[0025] The following embodiments will help those skilled in the art to further understand the present invention, but do not limit the present invention in any way. It should be noted that those skilled in the art can make several changes and improvements without departing from the concept of the present invention, and these all fall within the protection scope of the present invention.

[0026] The main raw materials and reagents used in the following examples and comparative examples have the following sources and specifications. Reagents not specifically mentioned are all commercially available analytical grade or higher grade products.

[0027] Direct reduced iron is commercially available high-purity sponge iron with an iron mass fraction ≥94wt%, a chromium mass fraction ≤0.02wt%, a nickel mass fraction ≤0.02wt%, and a silicon mass fraction ≤0.05wt%.

[0028] Carbon structural steel scrap is commercially available premium low-carbon steel offcuts, with a carbon content between 0.15wt% and 0.25wt%, a silicon content ≤0.1wt%, and no chromium or nickel alloying elements.

[0029] Preparation Example 1: This preparation example provides a method for preparing a low-silicon aluminum zirconium iron cored wire, including the following steps: 420g of zirconium powder and 580g of iron powder were weighed and thoroughly mixed. The mixed powder was placed in a yttrium oxide crucible in a medium-frequency induction furnace and heated to 1580℃ under an argon protective atmosphere for complete melting, allowing the zirconium and iron to fully react and form an alloy liquid. The molten alloy liquid was cast into a mold and allowed to cool naturally to room temperature. The alloy ingot was then crushed and sieved using a mechanical crushing device to collect low-silicon aluminum-zirconium-iron alloy powder with a particle size between 60 and 80 mesh. The collected low-silicon aluminum-zirconium-iron alloy powder was used as the inner core powder and continuously wrapped with a low-carbon steel strip with a thickness of 0.4mm and a width of 55mm to produce a low-silicon aluminum-zirconium-iron cored wire. Testing showed that the silicon mass fraction in the core powder was <1.0wt% and the aluminum mass fraction was <0.5wt%.

[0030] Preparation Example 2: This preparation example provides a method for preparing a low-silicon aluminum zirconium iron cored wire, including the following steps: 400g of zirconium powder and 600g of iron powder were weighed and thoroughly mixed. The mixed powder was placed in a yttrium oxide crucible in a medium-frequency induction furnace and heated to 1550℃ under an argon protective atmosphere for complete melting, allowing the zirconium and iron to fully react and form an alloy liquid. The molten alloy liquid was cast into a mold and allowed to cool naturally to room temperature. The alloy ingot was then crushed and sieved using a mechanical crushing device to collect low-silicon aluminum-zirconium-iron alloy powder with a particle size between 60 and 80 mesh. The collected low-silicon aluminum-zirconium-iron alloy powder was used as the inner core powder and continuously wrapped with a low-carbon steel strip with a thickness of 0.4mm and a width of 55mm to produce a low-silicon aluminum-zirconium-iron cored wire. Testing showed that the silicon mass fraction in the core powder was <1.0wt% and the aluminum mass fraction was <0.5wt%.

[0031] Preparation Example 3: This preparation example provides a method for preparing a low-silicon aluminum zirconium iron cored wire, including the following steps: 450g of zirconium powder and 550g of iron powder were weighed and thoroughly mixed. The mixed powder was placed in a yttrium oxide crucible in a medium-frequency induction furnace and heated to 1600℃ under an argon protective atmosphere for complete melting, allowing the zirconium and iron to fully react and form an alloy liquid. The molten alloy liquid was cast into a mold and allowed to cool naturally to room temperature. The alloy ingot was then crushed and sieved using a mechanical crushing device to collect low-silicon aluminum-zirconium-iron alloy powder with a particle size between 60 and 80 mesh. The collected low-silicon aluminum-zirconium-iron alloy powder was used as the inner core powder and continuously wrapped with a low-carbon steel strip with a thickness of 0.4mm and a width of 55mm to produce a low-silicon aluminum-zirconium-iron cored wire. Testing showed that the silicon mass fraction in the core powder was <1.0wt% and the aluminum mass fraction was <0.5wt%. Example 1

[0032] This embodiment provides a method for preparing and heat-treating high-strength alloy steel for tunnel boring machine cutters, including the following steps: (1) Batching and primary refining: The target chemical composition of the final alloy steel is calculated and weighed according to the following mass percentages: C: 0.60wt%, Mn: 1.4wt%, Mo: 3.0wt%, W: 1.75wt%, V: 1.25wt%, Ti: 0.075wt%, Zr: 0.02wt%, B: 0.0045wt%, Si: 0.05wt%, Ni: 0.05wt%, Cr: 0.05wt%, S: 0.003wt%, P: 0.005wt%, with the balance being Fe. The raw materials are weighed and converted based on the smelting yield of each alloying element. Graphite carburizing agent is added to adjust the carbon content according to the target carbon content. The base material consists of 75wt% direct reduced iron and 25wt% carbon structural steel scrap. The base material is added to an electric arc furnace for melting, oxygen blowing for decarburization, slag formation for dephosphorization, and then tapped to a ladle.

[0033] (2) Refining and sequential feeding: The ladle was transferred to the LF furnace for white slag desulfurization, and ferromanganese, ferromolybdenum, ferrotungsten, and ferrovanadium were added for basic alloying. Trace amounts of metallic nickel and metallic chromium (or corresponding nickel-iron and ferrochrome) were added according to the target chemical composition, and the temperature of the molten steel was adjusted to 1635℃. Then it was transferred to the VD vacuum degassing furnace and deep degassing was carried out for 18 minutes under a vacuum of 50 Pa. Four minutes before the vacuum was broken, the free oxygen activity of the molten steel was confirmed to be 12 ppm by a constant oxygen probe. The low-silicon aluminum zirconium iron cored wire prepared in Preparation Example 1 was fed into the molten steel at a wire feeding speed of 2.5 m / s, and the bottom was blown with argon and weakly stirred and allowed to stand for 1.5 minutes. Then, titanium iron alloy was added through the vacuum hopper, and the bottom was blown with argon and weakly stirred and allowed to stand for 1 minute to make it completely melted. Finally, boron iron cored wire with boron iron alloy powder as the internal core powder was fed in, and the bottom was blown with argon and allowed to stand for 1 minute. After breaking the vacuum, the molten steel is molded at 1540℃ to obtain alloy steel ingots.

[0034] (3) Forging and annealing: The alloy steel ingot is placed in a soaking furnace and heated to 1165°C at a rate of 15°C / min, and held for 2.5 hours. Multi-directional forging is carried out using a hydraulic press in the temperature range of 900°C to 1050°C, with a total forging ratio of 7. After forging, the forging is air-cooled to room temperature. Then, the forging is placed in an annealing furnace, heated to 830°C and held for 4.5 hours. It is then slowly cooled in the furnace to 510°C at a cooling rate of 15°C / h before being removed from the furnace and air-cooled to form a tool blank.

[0035] (4) Two-stage solution treatment: The tool blank is placed in a vacuum heat treatment furnace and the vacuum is evacuated to 0.05 Pa. In the first stage, the temperature is raised to 1070°C at a heating rate of 12°C / min and held for 1.5 hours. In the second stage, heating is stopped and high-purity nitrogen is introduced to slowly lower the furnace temperature to 935°C at a cooling rate of 4°C / min and hold at this temperature for 35 minutes.

[0036] (5) Salt bath isothermal graded quenching: After completing the two-stage solution treatment, the blade blank is transferred from the vacuum furnace within 15 seconds and immersed in a mixed molten nitrate salt bath with a composition of 50wt% potassium nitrate and 50wt% sodium nitrite. The salt bath temperature is controlled at 285℃ and isothermally maintained for 1.5 hours. After removal, it is allowed to cool naturally to room temperature in the air.

[0037] (6) In-situ cryogenic treatment: After the blank is cooled to room temperature, it is transferred to a cryogenic chamber within 1 hour. Liquid nitrogen vaporization is used for cooling, and the temperature is reduced to -125℃ at a rate of 2.5℃ / min. The temperature is maintained for 4.5 hours, and then the blank is slowly and naturally warmed to room temperature.

[0038] (7) Two-stage targeted tempering: Place the cryogenic workpiece into the tempering furnace, heat it to 525℃ and hold it for 2.5 hours, then remove it from the furnace and air cool it to room temperature; heat it again to 555℃ and hold it for 2.5 hours, then remove it from the furnace and air cool it to room temperature. Example 2

[0039] This embodiment provides a method for preparing and heat-treating high-strength alloy steel for tunnel boring machine cutters, including the following steps: (1) Batching and primary refining: The target chemical composition of the final alloy steel is calculated and weighed according to the following mass percentages: C: 0.55wt%, Mn: 1.2wt%, Mo: 2.5wt%, W: 1.5wt%, V: 1.0wt%, Ti: 0.05wt%, Zr: 0.015wt%, B: 0.003wt%, Si: 0.03wt%, Ni: 0.02wt%, Cr: 0.02wt%, S: 0.002wt%, P: 0.003wt%, with the balance being Fe. The raw materials are weighed and converted based on the smelting yield of each alloying element. Graphite carburizing agent is added to adjust the carbon content according to the target carbon content. The base material consists of 80wt% direct reduced iron and 20wt% carbon structural steel scrap. The base material is added to an electric arc furnace for melting, oxygen blowing for decarburization, slag formation for dephosphorization, and then tapped into a ladle.

[0040] (2) Refining and sequential feeding: The ladle was transferred to an LF furnace for white slag desulfurization and basic alloying was performed by adding corresponding alloys, including manganese, molybdenum, tungsten, vanadium, and trace amounts of nickel and chromium raw materials calculated according to the target chemical composition. The temperature of the molten steel was adjusted to 1620℃. It was then transferred to a VD vacuum degassing furnace and deep degassed for 15 minutes under a vacuum of 60 Pa. Three minutes before breaking the vacuum, the free oxygen activity of the molten steel was confirmed to be 15 ppm. The low-silicon aluminum zirconium iron cored wire prepared in Preparation Example 2 was fed into the molten steel at a wire feeding speed of 2.0 m / s, and the bottom was blown with argon and weakly stirred and allowed to stand for 1 minute; then titanium iron alloy was added, and the bottom was blown with argon and weakly stirred and allowed to stand for 1 minute; finally, boron iron cored wire was fed in, and the bottom was blown with argon and allowed to stand for 1 minute. After breaking the vacuum, the alloy steel ingot was obtained by die casting at 1530℃.

[0041] (3) Forging and annealing: Heat the alloy steel ingot to 1150℃ and hold for 2 hours. Perform multi-directional forging in the temperature range of 900℃ to 1050℃, with a total forging ratio of 6. After forging, air cool to room temperature. Heat the forging to 820℃ and hold for 4 hours. Cool it slowly in the furnace at a cooling rate of 15℃ / h to 500℃ and then air cool it to form a tool blank.

[0042] (4) Two-stage solution treatment: The tool blank is placed in a vacuum heat treatment furnace and the vacuum is evacuated to 0.08 Pa. In the first stage, the temperature is heated to 1060℃ at a heating rate of 10℃ / min and held for 1.5 hours. In the second stage, high-purity nitrogen is introduced to slowly reduce the furnace temperature to 920℃ at a cooling rate of 3℃ / min and hold at this temperature for 30 minutes.

[0043] (5) Salt bath isothermal graded quenching: Within 15 seconds, the blade blank is transferred and immersed in a 280℃ nitrate salt bath with a composition of 50wt% potassium nitrate and 50wt% sodium nitrite. It is kept isothermally for 1.5 hours, and then air-cooled to room temperature.

[0044] (6) In-situ cryogenic treatment: Transfer to a cryogenic chamber within 1 hour, cool to -120℃ at a cooling rate of 2.0℃ / min, keep warm for 4 hours, and slowly return to room temperature with the chamber.

[0045] (7) Two-stage targeted tempering: heat to 520℃ and hold for 2 hours, then remove from the furnace and air cool to room temperature; heat again to 550℃ and hold for 2 hours, then remove from the furnace and air cool to room temperature. Example 3

[0046] This embodiment provides a method for preparing and heat-treating high-strength alloy steel for tunnel boring machine cutters, including the following steps: (1) Batching and primary refining: The target chemical composition of the final alloy steel is calculated and weighed according to the following mass percentages: C: 0.65wt%, Mn: 1.6wt%, Mo: 3.5wt%, W: 2.0wt%, V: 1.5wt%, Ti: 0.1wt%, Zr: 0.025wt%, B: 0.006wt%, Si: 0.08wt%, Ni: 0.08wt%, Cr: 0.08wt%, S: 0.004wt%, P: 0.008wt%, with the balance being Fe. The raw materials are weighed and converted based on the smelting yield of each alloying element. Graphite carburizing agent is added to adjust the carbon content according to the target carbon content. The base material consists of 70wt% direct reduced iron and 30wt% carbon structural steel scrap. The base material is added to an electric arc furnace for melting, oxygen blowing for decarburization, slag formation for dephosphorization, and then tapped into a ladle.

[0047] (2) Refining and sequential feeding: The ladle was transferred to an LF furnace for white slag desulfurization and basic alloying, including the addition of manganese, molybdenum, tungsten, vanadium, and trace amounts of nickel and chromium raw materials calculated according to the target chemical composition, and the temperature was adjusted to 1650℃. It was then transferred to a VD vacuum degassing furnace and maintained at a vacuum of 40 Pa for 20 minutes for deep degassing. Five minutes before breaking the vacuum, the free oxygen activity was confirmed to be 10 ppm. The low-silicon aluminum zirconium iron cored wire prepared in Preparation Example 3 was fed in at a speed of 3.0 m / s, and the bottom was blown with argon and weakly stirred and allowed to stand for 2 minutes; then titanium iron alloy was added, and the bottom was blown with argon and weakly stirred and allowed to stand for 1 minute; finally, boron iron cored wire was fed in, and the bottom was blown with argon and allowed to stand for 1 minute. After breaking the vacuum, the steel ingot was obtained by die casting at 1550℃.

[0048] (3) Forging and annealing: Heat the alloy steel ingot to 1180℃ and hold for 3 hours. Perform multi-directional forging at 900℃ to 1050℃ with a total forging ratio of 8, and air cool to room temperature. Heat the forging to 840℃ and hold for 5 hours, then slowly cool it to 520℃ at a cooling rate of 20℃ / h before air cooling to form a preliminary blank.

[0049] (4) Two-stage solution treatment: The tool blank is placed in a vacuum furnace and evacuated to 0.02 Pa. In the first stage, it is heated to 1080°C at 15°C / min and held for 2 hours. In the second stage, high-purity nitrogen is introduced to cool the furnace temperature to 950°C at a cooling rate of 5°C / min and held for 40 minutes.

[0050] (5) Salt bath isothermal graded quenching: Immerse in a 290℃ nitrate salt bath with 50wt% potassium nitrate and 50wt% sodium nitrite for 15 seconds, maintain isothermal for 2 hours, and air cool to room temperature.

[0051] (6) In-situ cryogenic treatment: Transfer to a cryogenic chamber within 1 hour, cool down to -130℃ at 3.0℃ / min, keep warm for 5 hours, and then return to room temperature.

[0052] (7) Two-stage targeted tempering: heat to 530℃ and hold for 3 hours, then air cool to room temperature; heat again to 560℃ and hold for 3 hours, then air cool to room temperature.

[0053] Comparative Example 1: Compared with Example 1, the difference lies in the formulation and heat treatment process. This comparative example uses a conventional high-carbon, high-chromium D2 tool steel formulation (carbon mass fraction of 1.5wt%, chromium mass fraction of 12wt%, without the addition of tungsten, zirconium, and boron microalloying elements), and the heat treatment process is conventional austenitization at 1030℃ followed by direct oil quenching, and then conventional double-heat tempering at 520℃. No double-stage solution cooling, salt bath isothermal quenching, or deep cryogenic treatment is performed. All other processes are the same.

[0054] Comparative Example 2: Compared with Example 1, the difference is that zirconium and boron are not added to the alloy composition formulation (the Zr and B content in the target formulation is 0, and the balance is made up by Fe), and the step of adding low-silicon aluminum zirconium iron cored wire and boron iron cored wire is omitted during the refining period, while the rest are the same.

[0055] Comparative Example 3: Compared with Example 1, the difference lies in the order of feeding at the end of refining. In this comparative example, after confirming that the free oxygen activity of the molten steel meets the standard, the low silicon aluminum zirconium iron cored wire, titanium iron alloy and boron iron cored wire are added to the molten steel at the same time. After bottom argon blowing and stirring, the mixture is directly poured into the molten steel after standing. The time-sequential step feeding control is cancelled. All other aspects are the same.

[0056] Comparative Example 4: Compared with Example 1, the difference is that the second stage of the two-stage solution treatment is cancelled (that is, the cooling micro-segregation stage of introducing nitrogen to cool down to 935°C and holding for 35 minutes is cancelled). After the tool blank is homogenized at 1070°C for 1.5 hours in the first stage, it is directly transferred and immersed in a nitrate bath at 285°C for isothermal quenching. All other aspects are the same.

[0057] Comparative Example 5: Compared with Example 1, the difference is that the in-situ deep cryogenic treatment step is cancelled. After the blade blank completes the isothermal graded quenching in salt bath and is naturally cooled to room temperature in air, it is directly placed into the tempering furnace for two-stage targeted tempering. All other aspects are the same.

[0058] Test Example 1: The experimental steps are as follows: (1) Prepare standard cylindrical specimens for thermal expansion testing. The specimens are 4 mm in diameter and 10 mm in length. The specimens are taken from the forged annealed blanks of Examples 1, 2, 3 and Comparative Examples 4 and 5 before undergoing two-stage solution treatment (Comparative Example 4 is a single-stage solution treatment).

[0059] (2) The above sample was placed into the heating chamber of a high-precision quenching thermal expansion instrument and thermal simulation was performed using the same heat treatment process parameters as those in each embodiment and comparative example.

[0060] (3) Record the relative expansion data of the sample as it cools to the isothermal quenching temperature and remains isothermal for a period of time. Calculate the time from the start of isothermal treatment to when the expansion curve deviates significantly from the baseline as the phase transformation incubation period, and extract the maximum relative expansion during the entire isothermal stage.

[0061] (4) The sample after the simulated heat treatment was ground and polished, and the phase was determined by X-ray diffraction. The test used a Cu-Kα radiation source, the tube voltage was set to 40kV, the tube current was set to 40mA, the scanning range was set to 2θ=35° to 105°, and the scanning step was 0.02°.

[0062] (5) Obtain diffraction pattern data, use Rietveld full spectrum fitting method and direct comparison method to extract the integral intensity of characteristic diffraction peaks of face-centered cubic austenite phase and calculate the volume fraction of residual austenite; at the same time, combine the metallographic / scanning electron microscope images after etching to perform statistical analysis on the volume fraction of lower bainite in the final microstructure of the sample.

[0063] The experimental results are shown in Table 1.

[0064] Table 1. Test data on bainitic phase transformation kinetics and phase composition in the examples and comparative examples: Example 1 148 4.23 85.6 2.4 Example 2 162 4.05 82.1 3.1 Example 3 135 4.41 88.3 1.8 Comparative Example 4 1245 1.56 31.4 18.7 Comparative Example 5 151 4.21 84.5 11.2 From Table 1 and in combination Figure 1The results show that Examples 1-3 differ from Comparative Examples 4 and 5 in terms of isothermal transformation behavior and final phase composition. In Examples 1-3, under the condition of two-stage solution cooling treatment, the phase transformation incubation period was in the range of 135s-162s, the maximum relative expansion during the isothermal stage was relatively high, and the volume fraction of lower bainite in the final microstructure was all above 80 vol%, while the volume fraction of retained austenite was relatively low.

[0065] Compared to Examples 1-3, Comparative Example 4, which omits the second-stage cooling and holding treatment, exhibits a significantly prolonged phase transformation incubation period, a lower maximum relative expansion during the isothermal stage, and a significantly decreased lower bainite volume fraction in the final microstructure, while the retained austenite volume fraction increases accordingly. These results suggest that the second-stage cooling and holding process may help regulate the distribution of alloying elements in austenite and, to some extent, reduce the stability of supercooled austenite, thereby facilitating the subsequent bainite transformation during the isothermal stage.

[0066] Furthermore, compared to Comparative Example 5 (which omits cryogenic treatment), Example 1 showed little difference in the phase transformation incubation period and maximum relative expansion during the isothermal phase. However, the final retained austenite volume fraction in Comparative Example 5 was significantly higher than that in Example 1. This indicates that cryogenic treatment may have a positive effect on reducing the final retained austenite content and contribute to improving microstructure stability.

[0067] Summary Table 1 and Figure 1 The results show that the combined process of two-stage solution cooling treatment and subsequent cryogenic treatment described in this application has a certain promoting effect on controlling the final microstructure, increasing the amount of lower bainite transformation, and reducing the content of retained austenite.

[0068] Test Example 2: The experimental steps are as follows: (1) Prepare electrolytic physical extraction samples. The samples are taken from the final finished alloys of Examples 1-3 and Comparative Examples 1 and 3 after undergoing a complete heat treatment process (including tempering). The samples are processed into cylinders with a diameter of 10 mm and a length of 20 mm, and their surfaces are mechanically ground and polished to remove the oxide layer.

[0069] (2) The treated sample was used as the anode, and a platinum mesh was used as the cathode. The sample was placed in an electrolytic cell for constant current density electrolytic extraction. The electrolyte was a 10% (v / v) hydrochloric acid-methanol solution. The electrolysis process was carried out in a cryogenic constant temperature bath below -5°C, with the current density set at 0.05 A / cm². 2 Electrolysis continues until the matrix on the sample surface is completely dissolved and the internal precipitated phase is released.

[0070] (3) After electrolysis, the bottom sediment of the electrolyte containing the precipitated phase is collected and vacuum filtered using a microporous membrane with a pore size of 0.1 μm. The filter residue is repeatedly washed with anhydrous ethanol to remove residual iron salts and then dried in a vacuum drying oven at room temperature for 24 hours to obtain pure carbide / nitride precipitated phase extract powder.

[0071] (4) The dried extraction powder was uniformly coated on the quartz glass sample stage and phase scanning was performed using an X-ray diffractometer. A Cu-Kα radiation source was used, the scanning angle 2θ range was set to 30° to 80°, and the scanning rate was 1° / min. A precise diffraction pattern was obtained. The main characteristic peaks were calibrated by searching the standard PDF powder diffraction card, and the integrated intensity value of the characteristic diffraction peaks of specific crystal planes was recorded to characterize the relative content of the phase.

[0072] The experimental results are shown in Table 2.

[0073] Table 2. Integrated intensity test data of X-ray diffraction characteristic peaks of the extract phases in the examples and comparative examples: Example 1 425.6 681.3 Not detected 512.8 21.4 Example 2 398.2 653.1 Not detected 487.6 18.7 Example 3 467.5 712.9 Not detected 545.2 26.3 Comparative Example 1 Not detected Not detected Not detected Not detected 1856.4 Comparative Example 3 42.1 185.4 563.7 498.3 24.5 From Table 2 and in combination Figure 2 The results show that the composition of the precipitated phases obtained after electrolytic physical extraction varies among different samples. In the extract phases of Examples 1-3, relatively obvious characteristic peaks of ZrN, TiC, and M2C were detected, while TiN was not detected or was below the detection limit; meanwhile, the intensity of the M7C3 characteristic peak was relatively low. These results indicate that, under the process conditions described in this application, the types of precipitated phases formed by elements such as zirconium, titanium, and molybdenum / tungsten are significantly different from those in traditional high-carbon, high-chromium systems.

[0074] Compared with Examples 1-3, the intensity of the ZrN characteristic peak in Comparative Example 3 was significantly reduced, the intensity of the TiC characteristic peak was also decreased, while the TiN characteristic peak appeared clearly. This result indicates that, under the condition of changing the feeding sequence at the end of refining, the reaction pathway between zirconium, titanium and nitrogen and carbon may change, thereby affecting the composition and relative content of the final precipitated phase.

[0075] Furthermore, no characteristic peaks of ZrN, TiC, TiN, and M2C were detected in Comparative Example 1, while the intensity of the characteristic peak of M7C3 was significantly increased. This indicates that the precipitate system of Comparative Example 1 differs significantly from that of the examples, and its precipitate type is mainly high-chromium carbides represented by M7C3.

[0076] Summary Table 2 and Figure 2The results show that the embodiments of this application exhibit a characteristic of coexistence of ZrN, TiC, and M2C in the precipitate composition, with a relatively low TiN content. In contrast, Comparative Example 3 and Comparative Example 1 show different trends, with TiN increasing and M7C3 becoming dominant, respectively. This difference, to some extent, suggests that the types of microalloying elements and their addition sequence may influence the final precipitate structure.

[0077] Test Example 3: The experimental steps are as follows: (1) The finished alloys of Examples 1-3 and Comparative Examples 1-5 after the complete heat treatment process were selected as experimental objects, and the samples were machined and surface treated in accordance with the mechanical property test specifications.

[0078] (2) Some samples were processed into hardness test blocks with dimensions of 10mm×10mm×20mm. The upper and lower surfaces were ground to be parallel and the surface roughness Ra was less than 0.8μm. The macroscopic overall hardness was tested at room temperature using a fully automatic Rockwell hardness tester, with the test scale set to HRC. Five different locations were randomly selected on the surface of each sample for indentation testing. The test data were recorded, and the highest and lowest values ​​were discarded. The arithmetic mean of the remaining data was taken as the final macroscopic hardness value of the sample.

[0079] (3) The remaining specimens were processed into standard Charpy V-notch impact specimens with dimensions of 10mm × 10mm × 55mm, a notch depth of 2mm, and a notch bottom radius of curvature of 0.25mm. Impact fracture tests were performed on the specimens using a pendulum impact testing machine at room temperature of 20±5℃. The impact absorption energy Akv consumed by the pendulum to break the specimen was recorded. Three parallel specimens were prepared for each example and comparative example for repeated testing, and the arithmetic mean was taken to evaluate the impact fracture toughness of each alloy-steel material combination.

[0080] The experimental results are shown in Table 3.

[0081] Example 1 58.2 45.3 Example 2 57.5 47.1 Example 3 59.1 42.8 Comparative Example 1 60.5 8.4 Comparative Example 2 52.3 15.6 Comparative Example 3 56.4 12.1 Comparative Example 4 54.8 22.4 Comparative Example 5 55.6 25.7 As can be seen from the data in Table 3, there are significant differences between Examples 1-3 and Comparative Examples 1-5 in terms of macroscopic hardness and room temperature impact energy. The macroscopic Rockwell hardness of Examples 1 to 3 is all in the range of 57 HRC to 59 HRC, and the room temperature impact energy is all above 40 J, showing a good balance between strength and toughness.

[0082] In contrast, while Comparative Example 1 exhibited higher macroscopic hardness, its room-temperature impact absorption energy was significantly lower, indicating relatively insufficient toughness. Comparative Examples 2 and 3 both had lower hardness and impact absorption energy than Examples 1-3, with Comparative Example 3 showing a more significant decrease in impact absorption energy. These results suggest that the overall mechanical properties of materials may be adversely affected by the absence of appropriate microalloying elements or changes in the feeding sequence.

[0083] Furthermore, the macroscopic hardness of Comparative Example 4 and Comparative Example 5 were 54.8 HRC and 55.6 HRC, respectively, and their room temperature impact absorption energy were 22.4 J and 25.7 J, respectively, both lower than those of Examples 1-3. Considering the phase composition differences shown in Test Example 1, it can be concluded that the two-stage solution cooling treatment and cryogenic treatment may have a certain impact on the final microstructure and overall mechanical properties.

[0084] As can be seen from the results in Table 3, the embodiments of this application can still obtain relatively high room temperature impact absorption energy while maintaining a high hardness level, indicating that the combination of composition design and heat treatment process helps to improve the comprehensive mechanical properties of the material.

[0085] Test Example 4: The experimental steps are as follows: (1) Cut friction and wear test specimens from the finished specimens of Examples 1-3, Comparative Examples 1 and 4 after complete heat treatment, process them into cylindrical pin specimens with a diameter of 5 mm and a length of 12 mm, and grind the friction end face step by step until the surface roughness Ra reaches 0.2 μm to eliminate the influence of machining tool marks on the initial break-in period.

[0086] (2) All pin samples were thoroughly degreased and cleaned in acetone solvent for 15 minutes using an ultrasonic cleaner, and then dried and cooled to room temperature. The initial mass of each sample was weighed and recorded using a high-precision electronic analytical balance with a sensitivity of 0.1 mg.

[0087] (3) The pin sample was clamped on the pin-disc friction and wear testing machine. The grinding disc was a GCr15 bearing steel disc that had been hardened by quenching, and its surface hardness was set to 62HRC. The test parameters were set to dry sliding friction at room temperature without lubrication, with a normal load of 150N, a constant sliding linear velocity of 0.5m / s, and a total test sliding distance of 10000m. During the test, the built-in sensor of the equipment recorded the fluctuation data of the relative friction coefficient between the sample and the grinding disc in real time.

[0088] (4) After the friction test, the pin sample is removed and placed in acetone for ultrasonic cleaning to remove the wear debris and peeling material attached to the surface. After drying, the final mass after wear is weighed using an electronic analytical balance. The mass difference before and after the test is calculated as the wear loss, and the wear rate is calculated based on the load and sliding distance.

[0089] The experimental results are shown in Table 4.

[0090] Example 1 15.2341 15.2126 21.5 1.83 Example 2 15.1985 15.1752 23.3 1.98 Example 3 15.3412 15.3204 20.8 1.77 Comparative Example 1 15.2874 15.2642 23.2 1.97 Comparative Example 4 15.1456 15.0894 56.2 4.78 According to Table 4 and in conjunction with Figure 3 The results show that Examples 1-3 differ from Comparative Examples 1 and 4 in terms of wear weight loss, wear rate, and friction coefficient evolution characteristics. The wear weight losses of Examples 1-3 were 21.5 mg, 23.3 mg, and 20.8 mg, respectively, corresponding to wear rates of 1.83 × 10⁻⁶ mg and 1.83 × 10⁻⁶ mg, respectively. -6 mm 3 / N·m, 1.98×10 -6 mm 3 / N·m and 1.77×10 - 6 mm 3 / N·m, which is generally similar to Comparative Example 1, and the values ​​of some embodiments are slightly lower than those of Comparative Example 1.

[0091] Combination Figure 3 The evolution trend of the friction coefficient shown indicates that after the initial break-in period, the friction coefficient of Example 1 remained relatively stable with minimal fluctuations. In contrast, Comparative Example 1, while having an overall friction coefficient level similar to Example 1, exhibited more pronounced fluctuations during friction. This result suggests that although both examples demonstrate good wear resistance, differences may still exist in the frictional behavior of their contact surfaces.

[0092] Furthermore, the wear loss and wear rate of Comparative Example 4 were significantly higher than those of Examples 1-3 and Comparative Example 1, and Figure 3 The friction coefficient shown is relatively high overall, and it shows an increasing trend with the increase of sliding distance. This indicates that under the process conditions corresponding to Comparative Example 4, the wear resistance of the material surface may decrease.

[0093] Summary Table 4 and Figure 3 The results show that the embodiments of this application exhibit low wear weight loss and wear rate under dry sliding friction conditions, while the friction coefficient changes relatively steadily, indicating that the combination of composition design and heat treatment process has a certain positive effect on improving the friction and wear performance of the material.

[0094] Test Example 5: The experimental steps are as follows: (1) Cross-sectional specimens for simulated service testing were prepared from the forging entities of Examples 1-3, Comparative Example 1, and Comparative Example 5, which had completed all heat treatment processes. The specimens were machined into wedge-shaped simulated tool cutting edge structures with specific geometric features by using slow wire EDM combined with precision grinding technology. The wedge angle was set to 60°, the radius of curvature at the tip of the cutting edge was controlled at 1.5 mm, and the surface roughness was ground to Ra 0.4 μm to eliminate the interference of machining microcracks on fatigue crack initiation.

[0095] (2) The wedge-shaped specimen is rigidly fixed in the lower jaw fixture of the electro-hydraulic servo material fatigue testing machine. A high-hardness tungsten carbide hard alloy indenter with a flat end is used as an impact counter and installed on the upper servo actuation cylinder. The initial clearance is adjusted and the testing machine is set to force control mode.

[0096] (3) A unidirectional pulsating alternating compressive load was applied in a room temperature air environment. The maximum impact peak load was set to 25kN, the load ratio R was 0.1, the loading waveform was a sine wave, and the loading frequency was set to 15Hz. During the test, the dynamic displacement sensor integrated into the equipment was used to record the evolution of the dynamic stiffness of the loading system synchronously. An external high-definition industrial camera was used to capture the macroscopic morphology of the cutting edge area at set time intervals.

[0097] (4) When the dynamic stiffness of the system suddenly decreases by a set threshold (15%), or when the industrial camera detects a macroscopic spalling pit with a depth exceeding 2 mm on the cutting edge, the system automatically stops loading. Record the critical impact cycle number when the specimen fails, and collect the detached chipped fragments for macroscopic fracture archiving. Three parallel specimens are selected for testing in each evaluation group, and discrete data points are recorded.

[0098] The experimental results are shown in Table 5.

[0099] Example 1 60 25 248532 261024 254871 254809 Example 2 60 25 239655 228412 245108 237725 Example 3 60 25 268415 257930 262334 262893 Comparative Example 1 60 25 18563 24105 15472 19380 Comparative Example 5 60 25 86734 94521 82145 87800 Based on Table 5 and in conjunction with Figure 4 The results show that Examples 1-3 differ significantly from Comparative Examples 1 and 5 in terms of simulated cutting edge impact service life. The average failure cycles of Examples 1 to 3 were 254,809, 237,725, and 262,893, respectively, all significantly higher than those of Comparative Examples 1 and 5, indicating that the examples have higher impact resistance under the test conditions.

[0100] The average failure cycle count of Comparative Example 1 was 19,380, significantly lower than that of Examples 1-3; the average failure cycle count of Comparative Example 5 was 87,800, higher than that of Comparative Example 1, but still significantly lower than that of Examples 1-3. These results indicate that the lifespan of the samples under alternating impact loads varies considerably depending on the composition design and heat treatment process.

[0101] Combination Figure 4 The evolution trend of dynamic stiffness of the system shown indicates that the dynamic stiffness of Example 1 changes relatively smoothly over a longer cyclic range, while Comparative Examples 1 and 5 show a significant decrease in dynamic stiffness at an earlier stage and reach the failure threshold earlier. This suggests that Example 1 has a relatively better ability to maintain structural stability during impact cycling.

[0102] Summary Table 5 and Figure 4 The results show that, under the same wedge angle and peak impact load conditions, the embodiments of this application exhibit a higher mean failure cycle count and a relatively gentle dynamic stiffness decay trend, indicating that the combination of composition design and heat treatment process has a certain positive effect on improving the service reliability of materials under high-frequency impact conditions.

[0103] 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 high-strength alloy steel for tunnel boring machine cutters, characterized in that, It consists of the following chemical components by mass percentage: C: 0.55~0.65wt%; Mn: 1.2~1.6wt%; Mo: 2.5~3.5wt%; W: 1.5~2.0wt% V: 1.0~1.5wt%; Ti: 0.05~0.1wt%; Zr: 0.015~0.025wt%; B: 0.003~0.006wt%; Si: 0.03~0.08wt% Ni: 0.02~0.08wt% Cr: 0.02~0.08wt%; S: ≤0.004wt% P: ≤0.008wt% The balance consists of Fe and unavoidable impurities; The alloy steel is made by degassing and refining molten steel containing the above-mentioned chemical components in a vacuum, then feeding low-silicon aluminum zirconium iron cored wire with low-silicon aluminum zirconium iron alloy powder as the core powder to deoxidize and reduce the free oxygen activity of the molten steel. Subsequently, titanium iron alloy and boron iron cored wire are added in sequence. After casting and forging, the steel is then subjected to two-stage solution treatment, isothermal graded quenching in salt bath, in-situ deep cryogenic treatment, and two-stage targeted tempering.

2. The high-strength alloy steel for tunnel boring machine cutters according to claim 1, characterized in that, The alloy steel is composed of the following chemical components by weight percentage: C: 0.60 wt%, Mn: 1.4 wt%, Mo: 3.0 wt%, W: 1.75 wt%, V: 1.25 wt%, Ti: 0.075 wt%, Zr: 0.02 wt%, B: 0.0045 wt%, Si: 0.05 wt%, Ni: 0.05 wt%, Cr: 0.05 wt%, S: 0.003 wt%, P: 0.005 wt%, with the balance being Fe and unavoidable impurities.

3. A heat treatment method for high-strength alloy steel for tunnel boring machine cutters as described in claim 1 or 2, characterized in that, Includes the following steps: (1) Batching and primary refining: Based on the target chemical composition, the raw materials are calculated and weighed, the bottom material is added to the electric arc furnace for melting, oxygen blowing for decarburization, slag formation for dephosphorization, and then the steel is tapped to the ladle; (2) Refining and sequential feeding: The ladle is transferred to the LF furnace for desulfurization and basic alloying, and then transferred to the VD vacuum degassing furnace for deep degassing; before breaking the vacuum, after confirming that the free oxygen activity of the molten steel meets the standard, low silicon aluminum zirconium iron cored wire, titanium iron alloy, and boron iron cored wire are fed into the molten steel in sequence, and alloy steel ingots are obtained by die casting after breaking the vacuum. (3) Forging and annealing: After heating and holding the alloy steel ingot, it is forged in multiple directions, then annealed and slowly cooled to form a tool blank; (4) Two-stage solution treatment: The tool blank is heated and kept at a certain temperature in a vacuum heat treatment furnace for the first stage of solution treatment; then the heating is stopped and high-purity nitrogen is introduced to cool down and keep it at a certain temperature for the second stage of solution treatment. (5) Salt bath isothermal graded quenching: After completing the two-stage solution treatment, the knife blank is quickly transferred and immersed in a mixed molten nitrate bath for isothermal holding. After taking it out, it is air-cooled to room temperature. (6) In-situ cryogenic treatment: The blank is transferred into a cryogenic chamber, cooled and kept warm with liquid nitrogen, and then slowly and naturally warmed to room temperature; (7) Two-stage targeted tempering: The shield cutter is made of high-strength alloy steel by going through two heating, holding and air cooling steps at different temperatures.

4. The heat treatment method according to claim 3, characterized in that, In step (1), the base material consists of 70-80 wt% direct reduced iron and 20-30 wt% carbon structural steel scrap; and during the initial smelting process, graphite carbonizer is added to adjust the carbon content according to the target carbon content.

5. The heat treatment method according to claim 3, characterized in that, The specific implementation method of step (2) is as follows: Adjust the temperature of the molten steel in the LF furnace to 1620–1650℃; Transfer to a VD vacuum degassing furnace and maintain a vacuum of 40–60 Pa for 15–20 minutes for deep degassing; 3-5 minutes before breaking the vacuum, confirm that the free oxygen activity of the molten steel is 10-15 ppm; feed low-silicon aluminum zirconium iron cored wire into the molten steel at a feeding speed of 2.0-3.0 m / s and gently stir with argon blowing from the bottom for 1-2 minutes; then add titanium iron alloy and gently stir with argon blowing from the bottom for 1 minute; finally feed boron iron cored wire with boron iron alloy powder as the internal core powder and gently stir with argon blowing from the bottom for 1 minute; after breaking the vacuum, perform die casting at 1530-1550℃.

6. The heat treatment method according to claim 3, characterized in that, In step (2), the preparation method of the low-silicon aluminum zirconium iron cored wire is as follows: Mix 40-45 wt% zirconium powder and 55-60 wt% iron powder, place them in a yttrium oxide crucible in a medium-frequency induction furnace, and heat them to 1550-1600℃ under an argon protective atmosphere to completely melt them into an alloy liquid; after casting and natural cooling, crush and sieve the low-silicon aluminum zirconium iron alloy powder with a particle size of 60-80 mesh. The collected low-silicon aluminum zirconium iron alloy powder is used as the inner core powder and is continuously wrapped and formed by a low-carbon steel strip with a thickness of 0.4 mm and a width of 55 mm; wherein, the silicon mass fraction in the core powder is <1.0 wt% and the aluminum mass fraction is <0.5 wt%.

7. The heat treatment method according to claim 3, characterized in that, In step (3), the control parameters are as follows: the alloy steel ingot is heated to 1150-1180℃ at a rate of 15℃ / min and held for 2-3 hours; multi-directional forging is carried out in the temperature range of 900℃ to 1050℃, the total forging ratio is 6-8, and the ingot is air-cooled to room temperature after forging; the annealing heating temperature is 820-840℃, held for 4-5 hours, and then slowly cooled in the furnace at a cooling rate of 15-20℃ / h to 500-520℃ before being air-cooled.

8. The heat treatment method according to claim 3, characterized in that, The specific process parameters for step (4) are as follows: evacuate to 0.02-0.08 Pa; in the first stage, heat to 1060-1080°C at a heating rate of 10-15°C / min and hold for 1.5-2 hours; in the second stage, introduce high-purity nitrogen gas to slowly reduce the furnace temperature to 920-950°C at a cooling rate of 3-5°C / min and hold at this temperature for 30-40 minutes.

9. The heat treatment method according to claim 3, characterized in that, In step (5), the blade blank is transferred and immersed in a mixed molten nitrate bath containing 50wt% potassium nitrate and 50wt% sodium nitrite within 15 seconds. The bath temperature is controlled at 280-290℃ and isothermally maintained for 1.5-2 hours. In step (6), the blank is transferred into a cryogenic chamber within 1 hour and cooled to -120 to -130°C at a cooling rate of 2.0 to 3.0°C / min, and held at that temperature for 4 to 5 hours.

10. The heat treatment method according to claim 3, characterized in that, In step (7), the specific parameters for the two-stage targeted tempering are as follows: in the first stage, the temperature is heated to 520-530℃ and held for 2-3 hours, then removed from the furnace and air-cooled to room temperature; in the second stage, the temperature is raised again to 550-560℃ and held for 2-3 hours, then removed from the furnace and air-cooled to room temperature.

Images