Method for controlling titanium inclusion in large casting blank GCr15 bearing steel

Abstract

The invention relates to the technical field of metallurgy, in particular to a control method for titanium inclusion in large casting blank GCr15 bearing steel. According to the method, control over the end point C-Ti of the electric arc furnace can be well achieved, meanwhile, titanium removal, denitrification and deoxidation are taken into consideration, and the electric arc furnace adopts eccentric bottom tapping and steel remaining operation to prevent TiO2-containing oxidizing slag in a primary smelting furnace from entering a steel ladle; a clean steel ladle is adopted for tapping, and low-titanium iron alloy and a carburant are alloyed to reduce the amount of Ti added to molten steel in the tapping process; in addition, by means of low-temperature tapping, recarburization with a low-nitrogen recarburizing agent and molten steel deoxidation depth control, molten steel nitrogen increasing in the tapping and LF refining process can be reduced, deep deoxidation of molten steel through RH vacuum chamber aluminum adding, treatment time under the ultimate vacuum degree is kept, molten steel denitrification through RH vacuum treatment is enhanced, and the production efficiency of the whole technological process is improved.

Description

Technical Field

[0001] The invention relates to the technical field of metallurgy, and in particular to a method for controlling titanium inclusions in a large-scale ingot GCr15 bearing steel. Background Art

[0002] The main bearing of the shield machine is the core and key component of the shield machine cutter drive system, which determines the operating status and service life of the shield machine. The diameter of the cylindrical roller used in the main bearing of large and super-large shield machines can reach Φ80~140mm, and it is subjected to harsh working conditions such as alternating heavy loads and time-varying impacts during service. Therefore, as the main force-bearing part of the shield machine, the cylindrical roller is required to have high strength, high toughness, high wear resistance, high precision and other characteristics. Large cylindrical rollers of shield machines are usually made of high-carbon chromium series materials (GCr15, GCr15SiMn, GCr18Mo, REGCr15). In order to ensure that large-section cylindrical rollers have good strength, toughness and uniform structure, ultra-pure smelted steel liquid is needed to be cast into large-section ingots. High-temperature homogenization treatment, large deformation ratio forging and rolling rapid cooling, high-efficiency spheroidizing annealing and other processes are used to eliminate uneven structural defects caused by macro-segregation and carbide liquid precipitation, so as to achieve the "high cleanliness and high homogeneity" preparation of large cylindrical roller materials.

[0003] To achieve high cleanliness and high homogeneity of materials, we must first solve the ultra-purification of bearing steel smelting and the macro-segregation control of large ingot casting from the source. With the rapid improvement of special steel smelting equipment and production control level, especially the control level of total oxygen content and oxide inclusions in high-carbon chromium bearing steel, the physical quality of domestic leading bearing steel production enterprises has reached w (TO) 0.0005% in steel, the maximum oxide inclusion size is 18μm, and the water immersion ultrasonic non-destructive testing technology can detect non-metallic inclusions larger than 10μm.

[0004] Studies have shown that as the TO content decreases, the number of large oxide inclusions in steel that serve as fatigue sources decreases significantly. At the same time, the brittle non-oxide inclusions that were originally a minor factor become the main fatigue source. For example, when the TO content in steel is reduced to 6×10 -6 When the titanium inclusions begin to dominate the rolling contact fatigue fracture of bearing steel, the titanium inclusions in high carbon chromium bearing steel are precipitated in the solid-liquid two-phase region during the solidification process of the molten steel. Although their size is smaller than the deoxidized inclusions Al2O3 in the steel, their hardness is significantly higher than Al2O3, and their thermal expansion coefficient is smaller than that of the base metal. Therefore, their harm to material fatigue failure is far greater than that of Al2O3. Under the condition of large ingot casting, there is a lack of technical measures to effectively control the precipitation of titanium inclusions during solidification.

[0005] Based on this, the present invention provides a method for controlling titanium inclusions in GCr15 bearing steel to solve the above-mentioned technical problems! Summary of the invention

[0006] The purpose of the present invention is to provide a method for controlling titanium inclusions in GCr15 bearing steel, which solves the problem of lack of effective control of solidification precipitation of titanium inclusions under the condition of large ingot casting.

[0007] To achieve the above object, the present invention provides the following technical solutions:

[0008] The present invention provides a method for controlling titanium inclusions in GCr15 bearing steel, comprising the following steps:

[0009] S1. Electric arc furnace smelting: molten iron and scrap steel are put into an electric arc furnace for smelting, oxygen is blown into the molten steel for decarburization, dephosphorization, denitrification and detitanium removal; the final carbon content of the electric arc furnace is controlled at 0.08-0.10%, the tapping temperature is controlled at 1550-1580°C, and an eccentric bottom tapping-steel retention operation mode is adopted to prevent the primary furnace slag from entering the ladle;

[0010] S2. Weak deoxidation and alloying of steel: Use a new steel ladle of the same steel grade without slag to collect the molten steel, and when one-third of the molten steel is discharged, add ferroalloy and carburizer to perform weak deoxidation and alloying on the molten steel;

[0011] S3, LF refining: After the steel is tapped, it enters the LF refining furnace for refining, and FeSi and SiC powder are used for deoxidation to produce white slag;

[0012] S4, RH vacuum treatment: RH vacuum degree is pumped to 67Pa, and aluminum blocks are added to the molten steel through the vacuum chamber to make the aluminum content in the molten steel reach 0.025-0.03%;

[0013] S5. Pouring: The large ingot mold casting method is adopted for casting, and argon gas is used to seal and protect the pouring during the process.

[0014] The present invention is further configured as follows: in the step S1, the ratio of molten iron to molten steel is 40-50%.

[0015] The present invention is further configured as follows: in the step S2, the deoxidation alloying is performed by selecting low-titanium high-carbon ferrochrome, low-titanium high-purity ferrosilicon, low-titanium ferromanganese and a carburizer for deoxidation alloying.

[0016] The present invention is further configured as follows: in step S3, the aluminum content is maintained at 0.01-0.015% during LF refining, and the titanium content is maintained at 0.0011-0.0013% after LF refining.

[0017] The present invention is further configured as follows: in step S3, the content of aluminum oxide in the refined slag is controlled at 20-30%, the basicity is 2.5-3.0, and the content of titanium dioxide is controlled at less than 0.1%.

[0018] The present invention is further configured as follows: in step S4, the treatment time under the ultimate vacuum degree is not less than 25 minutes.

[0019] Compared with the prior art, the present invention has the following beneficial effects:

[0020] 1. The present invention can control the C-Ti at the end point of the electric arc furnace, and at the same time, take into account detitanium, denitrification and deoxidation. The electric arc furnace adopts an eccentric bottom steel tapping and steel retention operation to prevent the TiO2-containing slag from the primary furnace from entering the ladle; clean ladle steel tapping, low-titanium ferroalloy and carburizer alloying are adopted to reduce the Ti increase of molten steel during the steel tapping process; in addition, low-temperature steel tapping, low-nitrogen carburizer carburization and control of the deoxidation depth of the molten steel can reduce the nitrogen increase of the molten steel during the steel tapping process, RH vacuum chamber aluminum is added to deeply deoxidize the molten steel, the processing time is maintained under the ultimate vacuum degree, and the RH vacuum treatment is strengthened to denitrify the molten steel, thereby improving the production efficiency of the entire process.

[0021] 2. For large ingot casting, accurate control of raw material impurities and strict production organization and management, strict control of Ti and N content in molten steel, and delaying the precipitation of titanium inclusions during the solidification process are effective measures to control the size of titanium inclusions in GCr15 bearing steel products. BRIEF DESCRIPTION OF THE DRAWINGS

[0022] Figure 1 Schematic diagram of the effect of cooling rate and TiN precipitation on Ti and N concentrations at the solidification front;

[0023] Figure 2 Schematic diagram of Ti and N enrichment and TiN inclusion precipitation in the solid-liquid two-phase region at the solidification front;

[0024] Figure 3 Schematic diagram of the effect of Ti and N content on the precipitation of TiN inclusions at the solidification front;

[0025] Figure 4 The initial Ti and N contents of GCr15 affect Ti(C x N 1-x ) Schematic diagram of the effect of solidification fraction on precipitation;

[0026] Figure 5 The effect of initial Ti and N contents of GCr15 on the precipitation of Ti(C x N 1-x ) Schematic diagram of the influence of composition;

[0027] Figure 6 Schematic diagram of the effect of cooling rate on the size of TiN inclusions precipitated during solidification of GCr15;

[0028] Figure 7 Schematic diagram of the effect of Ti and N content in GCr15 bearing steel on the size of TiN precipitated during solidification;

[0029] Figure 8 This is a schematic diagram of the changes in Ti and N content in steel during GCr15 smelting;

[0030] Fig. 9 This is a typical large-grain titanium inclusion observed in the GCr15 sample under an optical microscope;

[0031] Fig.10 This is the shape of titanium inclusions observed under a scanning electron microscope;

[0032] Fig.11 It is a complex phase inclusion formed by titanium inclusion and deoxidation product in GCr15 bearing steel sample;

[0033] Fig.12 This is a line scan of the energy spectrum of titanium inclusions under a scanning electron microscope. DETAILED DESCRIPTION

[0034] The following will be combined with the embodiments of the present invention to clearly and completely describe the technical solutions in the embodiments of the present invention. Obviously, the described embodiments are only part of the embodiments of the present invention, not all of the embodiments. Based on the embodiments of the present invention, all other embodiments obtained by ordinary technicians in this field without creative work are within the scope of protection of the present invention.

[0035] Example 1: Thermodynamic analysis of TiN precipitation in GCr15

[0036] Due to the selective crystallization phenomenon during the solidification of molten steel, the solute elements gradually enrich in the liquid phase at the solidification front after solidification begins, forming the segregation behavior of the solute elements. When the actual activity product of Ti and N in the liquid phase in the solid-liquid two-phase region at the solidification front is greater than the Ti-N equilibrium activity product at the solidification front temperature, TiN inclusions will precipitate according to reaction formula (1).

[0037] [Ti]+[N]=TiN (s)

[0038] ΔG 0 =-291000+107.91T (1)

[0039] lg(a Ti ·a N ) = lgK TiN =5.673-15200.8 / T s-l (2)

[0040] In formula (2), K TiN is the equilibrium activity product of TiN inclusion precipitation and the equilibrium constant of TiN formation reaction (1) The reciprocal of T s-lis the temperature of the solid-liquid two-phase region at the solidification front, K. As the solidification process progresses, T s-l The temperature drop from the liquidus temperature to the solidus temperature can be expressed by formula (3):

[0041] T s-l =T0-(T0-T l )·[1-g(T l -T s )·(T0-T s ) -1 ] -1 (3)

[0042] In formula (3), T0 is the melting point of pure iron, 1809K; T l and Ts are the liquidus temperature and solidus temperature, respectively, in K; g is the solidification fraction of the molten steel. As the solidification fraction of the molten steel increases, the actual activity product Q of Ti and N enriched in the liquid phase in the solid-liquid two-phase region at the solidification front TiN It can be expressed as:

[0043] Q TiN =f Ti ·f N ·w(Ti)·w(N) (4)

[0044] In formula (4), f Ti and f N are Ti and N at the solidification front temperature (T s-l ), w(Ti) and w(N) are the actual mass percentages of the elements Ti and N in the liquid phase of the solid-liquid two-phase region at the solidification front, respectively, which can be expressed by formula (5):

[0045]

[0046] Φ=4α·(1+4α) -1 (6)

[0047]

[0048] In formulas (5) to (7), Φ is the reverse diffusion coefficient of solute element i; k i is the equilibrium distribution coefficient of solute element i in γ-Fe and molten steel; α is the Fourier number; is the diffusion coefficient of solute element i in γ-Fe, cm 2 ·s -1 ; λ is the secondary crystal arm spacing, cm, which is related to the initial carbon content of the molten steel w(C)0 and the solidification cooling rate Rc (see formula (8)), K·min-1; τ is the temperature of the molten steel from the solidification front temperature (T s-l ) is cooled to the local solidification time of the solidus temperature (Ts) (see formula (9)), s.

[0049]

[0050] Table 1 Typical chemical composition of GCr15 bearing steel (%)

[0051]

[0052] Taking GCr15 as an example, assuming the composition is as shown in Table 1, the solidus and liquidus temperatures of the bearing steel are calculated by Factsage 8.0 as T s =1544K, T l =1723.8K. Solidification front temperature (T s-l )The activity coefficients f of Ti and N under Ti and f N Calculate according to formula (10) and formula (11) respectively:

[0053] lg f Ti =(2557 / T s-l -0.365)lg f Ti(1873) (10)

[0054] lg f N =(3280 / T s-l -0.75)lg f N(1873) (11)

[0055] In formula (10) and formula (11), f Ti(1873) and f N(1873) are the activity coefficients of Ti and N at 1873K, respectively, which can be calculated according to equations (12) and (13):

[0056]

[0057] According to the steel liquid composition and the interaction coefficient of solute element i on Ti and N at 1873K and Substituting into equation (12) and equation (13) respectively, we can obtain the activity coefficient f at 1873K Ti(1873) and f N(1873) , and then substitute into equation (10) and equation (11) to calculate the solidification front temperature (T s-l )The activity coefficients f of Ti and N under Ti and f N It should be pointed out that w(i) in equations (12) and (13) is the actual concentration of solute element i in the liquid phase of the solid-liquid two-phase region at the solidification front, and should be calculated according to equation (5). The equilibrium distribution coefficient k of solute element i in equations (5) and (7) is i The diffusion coefficients of γ-Fe and γ-Fe are shown in Table 2.

[0058] Table 2 Equilibrium distribution coefficient k of each element in molten steel between austenite and liquid phase i and the diffusion coefficient in γ-Fe

[0059]

[0060]

[0061] When the molten steel solidifies to Q TiN ≥K TiN When TiN inclusions are precipitated in the solid-liquid two-phase region at the solidification front, TiN inclusions will precipitate. When TiN begins to precipitate, the Ti and N atoms in the liquid phase around the inclusions are consumed due to the continuous growth of TiN inclusions, thereby affecting the segregation behavior of the solute elements and ultimately affecting the growth of inclusions. For this reason, the influence of inclusion precipitation factors needs to be coupled into the solute segregation model. The coupled model assumes that the precipitated TiN inclusions are uniformly distributed in the liquid phase of the solid-liquid two-phase region. When TiN begins to precipitate, there is a solute equilibrium between the surface of the inclusion particles and the nearby molten steel, and the equilibrium expression is:

[0062]

[0063] Δw(Ti) / Δw(N)=M Ti / M N (15)

[0064] Wherein, Δw(Ti) and Δw(N) represent the concentration values ​​of Ti and N consumed by TiN precipitation, respectively; is the equilibrium constant of TiN formation reaction (1); M Ti and M N are the molar masses of Ti and N, g·mol -1 During the inclusion precipitation process, the concentration values ​​of Ti and N are in equilibrium, as shown in the following formula:

[0065]

[0066] In formula (16), w(Ti) e and w(N) e Represent the equilibrium concentration values ​​of Ti and N respectively.

[0067] When TiN inclusions have not yet precipitated, Ti and N are continuously distributed in the interdendritic liquid phase. According to the segregation model of formula (5), if the entire solidification process is divided into several steps, the concentrations of Ti and N elements in the micro-unit between the (n-1)th step and the nth step are calculated according to formula (17) and formula (18), respectively:

[0068]

[0069] In formulas (17) and (18), w(i) n-1 and w(i) n Respectively represent the actual concentration of Ti or N element calculated in the (n-1)th step and the nth step; w(i)0 represents the initial concentration of Ti or N in the molten steel; Φ n-1 and Φ n are the Φ values ​​calculated in the (n-1)th step and the nth step, respectively. Combining equations (17) and (18), w(i) n It can be further expressed as:

[0070]

[0071] It should be noted that in the TiN precipitation model, the actual concentration of Ti or N in the calculation unit at the (n-1) step should be the concentration w(i) when the TiN inclusion precipitation reaches equilibrium. e Therefore, after TiN begins to precipitate, the segregation concentration w(i) in step n is calculated in equation (19): n The equilibrium concentration w(i) obtained in step (n-1) needs to be e Replace w(i) n-1 .

[0072] Figure 1 The results of the solidification segregation coupled with the TiN inclusion precipitation model of the GCr15 solute element composition in Table 1 are shown in Table 1. The mass percentage concentration of Ti and N enriched in the liquid phase in the solid-liquid two-phase region at the solidification front gradually increases with the increase of solidification fraction. When the actual activity product (Q TiN ) is greater than the equilibrium activity product of TiN precipitation at the solidification front temperature (K TiN ), TiN inclusions began to precipitate at the time of 0.976 solidification fraction of the molten steel, and the concentration curve had an inflection point. As TiN nucleated and grew, the concentration of N in the liquid phase around the inclusions decreased slowly, while the concentration of Ti decreased rapidly. This process can be described by Figure 2 In addition, Figure 1 It can be seen that the cooling rate is from 5℃·min -1 Increase to 50℃·min -1 The cooling rate has little effect on the solidification segregation of N, which may be due to the high diffusion rate of N in the molten steel. Before the precipitation of TiN, the cooling rate has a certain effect on the solidification segregation of Ti, especially from 10℃·min -1 Increase to 50℃min -1 hour.

[0073] Figure 3The thermodynamics of TiN inclusions precipitated in the solid-liquid two-phase region at the solidification front of GCr15 steel liquid at different initial Ti and N contents calculated by the coupling model. When the initial N (or Ti) content of the steel liquid is constant, reducing the initial Ti (or N) content can delay the precipitation time of TiN inclusions at the solidification front; when the initial N content of the steel liquid is 0.0025%, the initial Ti content is reduced to 0.00065% to avoid the precipitation of TiN inclusions. It is still very difficult to reduce the Ti in GCr15 to 0.00065% under the current BOF or EAF smelting conditions, but the possibility of further reducing N still exists. When the initial Ti of the steel liquid is 0.0010%, the initial N content is reduced to 0.0017% to avoid the precipitation of TiN inclusions. Therefore, for the casting of large-tonnage die-cast steel ingots, it is necessary to control the precipitation of TiN inclusions during the solidification process. The most effective measure is to control the Ti and N content in the molten steel during the steelmaking stage. In particular, reducing the Ti content in the molten steel has a greater effect on delaying the precipitation of TiN, such as scraping off the oxidized slag after tapping the converter to make new low-TiO2 slag, using low-Ti ferrochromium alloying, using vacuum refining and providing protective casting, etc.

[0074] Example 2: Ti(C) in GCr15 x N 1-x )Precipitation Thermodynamic Analysis

[0075] In addition to the precipitation of TiN inclusions, Ti(C x N 1-x ) inclusions. Ti(C x N 1-x ) is a continuous solid solution of TiN and TiC, and x is the molar ratio of C or TiC in the continuous solid solution, which varies between 0 and 1. x N 1-x ) has a higher hardness than TiN, so it is more harmful to material fatigue failure. Taking GCr15 as an example, after solidification begins, the solute elements Ti, N, and C in the molten steel continue to enrich in the liquid phase of the solid-liquid two-phase zone at the solidification front. When Ti(C x N 1-x ) precipitation equilibrium activity product, Ti(C x N 1-x )Inclusions:

[0076] [Ti]+x[C]+(1-x)[N]=Ti(C x N 1-x ) (20)

[0077] In the continuous solid solution Ti(C x N 1-x ) formation reaction, the equilibrium concentration of inclusion-forming elements must satisfy the following relationship

[19] :

[0078]

[0079] In formula (21) and formula (22), w(Ti)e, w(C)e and w(N)e are the equilibrium concentration values ​​(%) of Ti, C and N in the molten steel at the solidification front temperature, respectively; and are the equilibrium concentration products of the TiC and TiN formation reactions, respectively, calculated according to equations (23) and (24).

[0080]

[0081] When the concentrations of Ti, N and C in the solid-liquid two-phase region at the solidification front are enriched to the point where their actual concentration product is greater than or equal to the equilibrium concentration product, the solidification front has the thermodynamic conditions for the precipitation of Ti(CxN1-x). Therefore, w(Ti)e, w(C)e and w(N)e in equations (21) and (22) should be replaced by w(Ti), w(C) and w(N), thus obtaining:

[0082]

[0083] The actual concentrations w(Ti), w(C) and w(N) of Ti, N and C at the solidification front of the molten steel are calculated according to formula (5). In order to simplify the calculation, let the inverse diffusion coefficient Φ in formula (5) = 0 or 1, and the non-equilibrium solidification equation (i.e., Scheil equation, formula (27)) and the equilibrium solidification equation (formula (28)) are obtained to calculate the solidification segregation of the solute element i. The former is equivalent to the solute element being completely non-diffused in the solid phase, while the latter is completely diffused in the solid phase:

[0084]

[0085] w(i)=w(i)0[1-(1-k i )g] -1 (28)

[0086] Formula (27) is used to calculate the actual concentration of Ti at the solidification front, and formula (28) is used to calculate the actual concentration of C and N at the solidification front. Substituting formula (27) and formula (28) into formula (25) and formula (26) respectively, we can obtain:

[0087]

[0088]

[0089] Combining equations (29), (30) and (3), we can calculate the Ti(C) precipitation during solidification of the GCr15 bearing steel with the composition in Table 1. x N 1-x ) inclusions, the solidification fraction g of the molten steel is 0.937, and the solidification front temperature Ts-l At 1302℃, the precipitated Ti(C x N 1-x ) inclusions has an x ​​value of 0.11. This indicates that during the solidification of GCr15 steel liquid, Ti(C 0.11 N 0.89 ) will precipitate before TiN inclusions at the solidification front.

[0090] Effect of initial Ti and N contents in molten steel on the precipitation of Ti(C x N 1-x ) when the temperature (T s-l ), solidification fraction (g) and composition (x) have a significant impact. Figure 4 and Figure 5 The solidification precipitation of Ti(C) in GCr15 steel liquid with different w(Ti)0 and w(N)0 are given respectively. x N 1-x ) inclusion solidification fraction (g) and composition (x). As the initial Ti and N content of the molten steel decreases, the Ti(C x N 1-x ) is included in the precipitation at the solidification front. At the same time, according to formula (3), Ti(C x N 1-x ) precipitation temperature also decreases. Reducing w(Ti)0 can make the precipitated Ti(C x N 1-x ) The x value in the inclusion decreases, but the decrease of w(N)0 makes Ti(C x N 1-x )The x value in the inclusion increases.

[0091] Example 3: Kinetic analysis of titanium inclusion precipitation in GCr15 steel

[0092] Taking TiN inclusion as an example, after the crystal nucleus is formed, driven by the N concentration gradient in the solid-liquid two-phase region, N atoms diffuse to the TiN surface, and TiN particles continue to grow through interfacial chemical reactions. According to Fick's first law, the N diffusion flux (JN) is proportional to its concentration gradient:

[0093]

[0094] In formula (31), J N is the molar diffusion flux of N, mol·m -2 ·s -1 ; is the diffusion coefficient of nitrogen in the molten steel, r is the radius of TiN, m; CN is the molar concentration of N in the steel liquid, mol·m -3 ; C N eis the molar concentration of N when the reaction with TiN reaches equilibrium at the solidification front temperature, mol·m -3 By converting the molar concentration CN of N into the mass percentage concentration w(N), equation (31) can be transformed into equation (32):

[0095]

[0096] In formula (32), ρ Fe is the density of the molten iron (7.07t·m -3 );M N is the molar mass of N (14 g / mol).

[0097] According to the mass balance during the growth of TiN particles, the diffusion flux of N (J N ) and the TiN particle radius (r) are established as follows:

[0098]

[0099] Where MTiN in (33) is the molar mass of TiN (61.9 g·mol -1 );ρ TiN is the density of TiN (5.394t·m -3 ). Taking the derivative of the volume of TiN with respect to time, we can get:

[0100]

[0101] Combining equations (32) to (34), the following equation can be derived:

[0102]

[0103] Integrating the above formula gives:

[0104]

[0105] In formula (36), w(N), w(N) e and τ are calculated according to equation (5), equation (16) and equation (9) respectively. Figure 6 The relationship between the size change of TiN inclusions precipitated during solidification and the solidification cooling rate of GCr15 bearing steel with the composition in Table 1 calculated by formula (36) is shown in Table 1. Figure 6 It can be seen that: (1) As the solidification of the molten steel progresses, the size of the TiN inclusion increases in a parabolic pattern. The inclusion growth rate is fast at the beginning, then the growth rate slows down, and finally reaches the maximum diameter at the end of solidification. (2) Although the cooling rate (RC) has little effect on the solidification segregation of Ti and N, it has a great influence on the growth of the TiN inclusion size. As the cooling rate increases from 5.0℃min -1 Increase to 10℃·min -1, the TiN inclusion size was reduced from 8.97μm to 6.33μm; the cooling rate was further increased to 50℃min -1 , TiN size decreased from 6.33μm to 2.96μm. However, for large ingot casting, it is technically difficult to control the solidification cooling rate.

[0106] Figure 7 The solidification cooling rate is shown as 10℃min -1 The influence of the initial Ti and N contents in GCr15 on the size of TiN inclusions precipitated during solidification is calculated. Figure 7 It can be seen that delaying the solidification and precipitation of TiN inclusions can reduce their final size. When the initial Ti (or N) content is constant, as the N (or Ti) content decreases, the solidification and precipitation time of TiN inclusions is delayed, and the final size becomes smaller. Figure 7 It can also be found that reducing the initial Ti content in the molten steel has a greater impact on the final size of TiN inclusions than reducing the initial N content in the molten steel. In other words, in actual production, the focus should be on how to further reduce the initial Ti content in the molten steel.

[0107] Example 4: Control of Ti in Molten Steel

[0108] S1. Electric arc furnace smelting: molten iron and scrap steel are mixed into molten steel and loaded into an electric arc furnace for smelting. Oxygen is blown into the molten steel for decarburization, dephosphorization, denitrification and detitanium. The final carbon content of the electric arc furnace is controlled at 0.08-0.1%, and the tapping temperature is controlled at 1550-1580°C. The eccentric bottom tapping-steel retention operation mode is adopted to prevent the primary furnace slag from entering the ladle.

[0109] Among them, the ratio of molten iron in the molten steel is 43%.

[0110] S2. Deoxidation and alloying of steel: Use a new steel ladle of the same steel grade without slag to collect the molten steel, and when one-third of the molten steel is discharged, add ferroalloy and carburizer for deoxidation and alloying.

[0111] Among them, deoxidation alloying is to use low titanium high carbon ferrochrome, low titanium high purity ferrosilicon, low titanium ferromanganese and carburizer for deoxidation alloying.

[0112] S3, LF refining: After the steel is tapped, it enters the LF refining furnace for refining, and FeSi and SiC powder are used for deoxidation to produce white slag.

[0113] The aluminum content is maintained at 0.01-0.015% during the refining process, and the titanium content is maintained at 0.0011-0.0013% after the refining is completed.

[0114] In the refined slag composition, the content of alumina is controlled at 20-30%, the basicity is controlled at 2.5-3.0, and the content of titanium dioxide is controlled at less than 0.1%.

[0115] S4. RH vacuum treatment: The RH vacuum degree is evacuated to 67Pa, and aluminum blocks are added to the molten steel through the vacuum chamber to make the aluminum content in the molten steel reach 0.025-0.03%.

[0116] Among them, the treatment time under the ultimate vacuum degree is not less than 25 minutes.

[0117] S5. Pouring: The large ingot mold casting method is adopted for casting, and argon gas is used to seal and protect the pouring during the process.

[0118] The change of Ti content in each process of GCr15 smelting is shown in Figure 8 .

[0119] Example 5: Control of N in Molten Steel

[0120] 1) Control of molten steel N in electric arc furnace steelmaking process

[0121] The nitrogen content in the high-quality scrap steel used in the electric arc furnace to smelt GCr15 bearing steel is relatively stable, while the w(N) in the molten iron varies between 0.0028% and 0.0062%. After the scrap steel is melted, the w(N) in the molten steel varies between 0.0050% and 0.0063%. The oxygen blowing decarburization process produces a large number of CO bubbles that escape from the molten pool and take away the nitrogen in the molten steel. At this time, the denitrification rate is proportional to the decarburization rate. The arc furnace end point w(C) is controlled at 0.08-0.10%, and a relatively high final decarburization rate can still be maintained to avoid nitrogen absorption by molten steel. The iron-to-steel ratio is 43%, and the end point temperature is lower than 1580°C, and the arc furnace end point w(N) can reach 0.0018-0.0022%.

[0122] 2) Control of N in molten steel during deoxidation and alloying

[0123] The nitrogen increase in molten steel during the deoxidation and alloying process mainly comes from nitrogen absorbed from the atmosphere, nitrogen brought in by alloys and recarburizers. Lower tapping temperatures and weak deoxidation of molten steel can reduce the amount of nitrogen absorbed by molten steel in the ladle from the atmosphere. The tapping temperature of an electric arc furnace is lower than that of a converter. At the same time, the addition of aluminum during tapping from an electric arc furnace is reduced in actual operation, which is beneficial to reducing nitrogen absorption by molten steel and also beneficial to LF refining to reduce the addition of titanium to molten steel from TiO2 in aluminum-reducing slag in the molten steel. The nitrogen brought in by the recarburizer during the alloying process accounts for the highest proportion. Graphite electrode powder with low nitrogen content or high-temperature electric calcined anthracite can be selected for recarburization. The w(N) of molten steel sampled at the LF station is between 0.0028 and 0.0033%.

[0124] 3) Molten steel N control during LF refining process

[0125] It is a common phenomenon that the molten steel increases nitrogen during the LF furnace refining process. Although the nitrogen capacity of slag systems of different compositions varies greatly, for the commonly used CaO-MgO-Al2O3-SiO2 slag system, no matter how high the partial pressure of nitrogen in the atmosphere is, as long as the molten steel surface is well covered by the slag layer, the nitrogen in the atmosphere cannot directly pass through the slag layer into the molten steel. Therefore, the exposure of molten steel during LF refining is the direct cause of the increase in nitrogen in the molten steel. Adding low-melting-point and high-basicity pre-melted slag to the ladle during the steel-tapping process can form a well-covered ladle top slag in advance, and maintain a sufficient slag layer thickness to prevent the arc from directly contacting the molten steel, which is beneficial to reduce nitrogen absorption in the early stage of refining. Controlling the argon blowing intensity during the LF refining process can prevent the molten steel from being blown bare, especially preventing the arc under the electrode from being blown bare. In actual production, the average nitrogen increase in the LF refining stage is 0.00085%.

[0126] 4)RH vacuum refining denitrification

[0127] Dissolved oxygen and sulfur in molten steel are surface active elements that hinder denitrification under vacuum. For deeply deoxidized and desulfurized molten steel, RH vacuum refining has excellent denitrification kinetic conditions. Under the ultimate vacuum degree, the denitrification rate of molten steel is controlled by the diffusion of the liquid phase boundary layer. Under the strong drive of the bubble pump in the RH riser, the diffusion resistance of the liquid phase boundary layer decreases, and the molten steel enters the vacuum chamber in batches and is directly exposed to the vacuum. Under extremely low nitrogen partial pressure, nitrogen atoms escape from the free surface of the molten steel and are sucked away by the vacuum; on the other hand, the nitrogen partial pressure in the argon bubbles in the riser is even lower, which is very conducive to the escape of nitrogen atoms in the molten steel. In actual production, the average denitrification of GCr15 after RH vacuum treatment reaches 0.0019%.

[0128] The change of nitrogen content in each process of GCr15 bearing steel smelting process is as follows Figure 8 shown.

[0129] Performance test: 60t ultra-pure GCr15 molten steel was prepared, and two 30t steel ingots were cast by pouring method under Ar gas protection, and hot forged / rolled into φ75mm round bars. The chemical composition of the samples was analyzed and shown in Table 3. As can be seen from Table 3, the TO, Ti and N contents of the finished product reached 0.0007%, 0.0010% and 0.0019% respectively, and the product has extremely high purity:

[0130] Table 3 Chemical composition of GCr15 bearing steel finished product (%)

[0131]

[0132] A 15mm×15mm×10mm sample was cut from the core of the round rod by wire cutting and ground into a metallographic sample to detect titanium inclusions in the steel. The detection surface was the cross section of the round rod. First, the titanium inclusions in the sample were observed under an OLYMPUSDSX1000 super depth of field / 3D microscope with a magnification of 2000 times. Fig. 9As shown in the figure, the titanium inclusions detected are all earthy red in color, and no golden yellow TiN inclusions are observed, indicating that the titanium inclusions in the 30t die-cast GCr15 bearing steel exist in the form of titanium carbonitride, which is consistent with the previous thermodynamic prediction. In terms of quantity, there are very few titanium carbonitride inclusion particles found in the 200 consecutively observed fields of view, but the proportion of those with a size greater than 10μm is relatively large, and the largest titanium inclusion size detected is 15μm (such as Fig. 9 (as shown in (a)).

[0133] The size distribution of titanium inclusions in GCr15 bearing steel samples was detected using ASIN EVO MA10 automatic inclusion analyzer, and the composition of typical titanium inclusions was scanned and analyzed. The typical titanium inclusion morphology observed under the scanning electron microscope is as follows: Fig.10 As shown in the figure, it can be found that the titanium inclusions in the die-cast GCr15 bearing steel have complex shapes, unlike the regular titanium inclusions observed in the hypereutectoid cord steel. In addition, some titanium carbonitride complex phase inclusions with deoxidation products MgO·Al2O3 or Al2O3 as the core were observed in the GCr15 bearing steel sample, as well as a small amount of titanium inclusions associated with Al2O3 or MgO·Al2O3 (see Fig.11 (c), (e)). The oxide inclusion particles that serve as the core of titanium inclusions are very small ( Fig.11 (a), (b)), from the energy spectrum composition, it is mainly MgO·Al2O3.

[0134] The element line scanning analysis of titanium inclusions was performed using an energy dispersive spectrometer under a scanning electron microscope. Fig.12 It can be found that the main elements in the titanium inclusions are Ti and N, but the inclusions obviously also contain C, indicating that the titanium inclusions precipitated during solidification in the GCr15 bearing steel are titanium carbonitride inclusions.

[0135] like Fig.10 The length of the long titanium inclusion shown in (a) reaches 16.2 μm. Fig.10 (h) also reached 17.4μm, and the average size of titanium inclusions detected by electron microscopy was 9.2μm. In general, the size of titanium inclusions detected in the die-cast GCr15 sample is much larger than that detected in the small billet continuous casting SWRH82A and SWRH92A in the past. From the perspective of inclusion growth dynamics, this is largely related to the low cooling rate of large-tonnage die-cast ingots. Therefore, for large ingot die casting, it is an effective measure to control the size of titanium inclusions in GCr15 bearing steel products through precise control of raw material impurities and careful production organization, strict control of Ti and N content in molten steel, and delaying the precipitation of titanium inclusions during solidification.

[0136] In the description of this specification, the description with reference to the terms "one embodiment", "example", "specific example", etc. means that the specific features, structures, materials or characteristics described in conjunction with the embodiment or example are included in at least one embodiment or example of the present invention. In this specification, the schematic representation of the above terms does not necessarily refer to the same embodiment or example. Moreover, the specific features, structures, materials or characteristics described can be combined in any one or more embodiments or examples in a suitable manner.

[0137] The preferred embodiments of the present invention disclosed above are only used to help illustrate the present invention. The preferred embodiments do not describe all the details in detail, nor do they limit the invention to the specific implementation methods described. Obviously, many modifications and changes can be made according to the content of this specification. This specification selects and specifically describes these embodiments in order to better explain the principles and practical applications of the present invention, so that those skilled in the art can understand and use the present invention well. The present invention is limited only by the claims and their full scope and equivalents.

Claims

1. A method for controlling titanium inclusions in large-scale ingot GCr15 bearing steel. It is characterized in that The following steps are involved: S1. Electric arc furnace smelting: molten iron and scrap steel are put into an electric arc furnace for smelting, oxygen is blown into the molten steel for decarburization, dephosphorization, denitrification and detitanium removal; the final carbon content of the electric arc furnace is controlled at 0.08-0.10%, the tapping temperature is controlled at 1550-1580°C, and an eccentric bottom tapping-steel retention operation mode is adopted to prevent the primary furnace slag from entering the ladle; S2. Weak deoxidation and alloying of steel: Use a new steel ladle of the same steel grade without slag to collect the molten steel, and when one-third of the molten steel is discharged, add ferroalloy and carburizer to perform weak deoxidation and alloying on the molten steel; S3, LF refining: After the steel is tapped, it enters the LF refining furnace for refining, and FeSi and SiC powder are used for deoxidation to produce white slag; S4, RH vacuum treatment: RH vacuum degree is pumped to 67Pa, and aluminum blocks are added to the molten steel through the vacuum chamber to make the aluminum content in the molten steel reach 0.025-0.03%; S5. Pouring: The large ingot mold casting method is adopted for casting, and argon gas sealing is used to protect the casting during the casting process.

2. A method for controlling titanium inclusions in a large-scale ingot GCr15 bearing steel according to claim 1, It is characterized in that In the step S1, the ratio of molten iron to molten steel is 40-50%.

3. The method for controlling titanium inclusions in a large-scale ingot GCr15 bearing steel according to claim 1, It is characterized in that In the step S2, the deoxidation alloying is performed by selecting low-titanium high-carbon ferrochrome, low-titanium high-purity ferrosilicon, low-titanium ferromanganese and a carburizer for deoxidation alloying.

4. A method for controlling titanium inclusions in a large-scale ingot GCr15 bearing steel according to claim 1, It is characterized in that In the step S3, the aluminum content is maintained at 0.01-0.015% during the LF refining process, and the titanium content is maintained at 0.0011-0.0013% after the LF refining is completed.

5. The method for controlling titanium inclusions in a large-scale ingot GCr15 bearing steel according to claim 1, It is characterized in that In step S3, the content of aluminum oxide in the refined slag is controlled to be 20-30%, the basicity is 2.5-3.0, and the content of titanium dioxide is controlled to be less than 0.1%.

6. A method for controlling titanium inclusions in large-scale ingot GCr15 bearing steel according to claim 1, It is characterized in that In step S4, the treatment time under the ultimate vacuum degree is not less than 25 minutes.

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