Nickel-based alloy

Abstract

To provide an alloy of a high γ' volume fraction having particularly excellent resistance to hot cracking.SOLUTION: A nickel-based alloy composition comprises 4.0 to 5.6 mass% of Al, 0.0 to 1.0 mass% of Ti, 0.0 to 4.0 mass% of Nb, 0.0 to 11.9 mass% of Ta, 2.0 to 12.7 mass% of W, 0.0 to 3.0 mass% of Mo, 0.0 to 22.0 mass% of Co, 6.0 to 16.7 mass% of Cr, 0.02 to 0.35 mass% of C, 0.001 to 0.2 mass% of B, 0.00 to 0.5 mass% of Zr, 0.0 to 3.0 mass% of Re, 0.0 to 2.0 mass% of Ru, 0.0 to 3.0 mass% of Ir, 0.0 to 0.5 mass% of V, 0.0 to 0.003 mass% of S, 0.0 to 0.25 mass% of Mn, 0.0 to 0.1 mass% of Mg, 0.0 to 4.0 mass% of Fe, 0.0 to 0.5 mass% of Cu, 0.0 to 1.0 mass% of Hf, the balance being Ni and incidental impurities, wherein the relationship among Nb, Ta, and W contained in the alloy satisfies a predetermined formula.SELECTED DRAWING: None

Description

Technical Field

[0001] The present invention relates to nickel-based superalloy compositions designed for application in additive manufacturing (AM) processes. Examples of such processes include, but are not limited to, powder bed-based AM methods (e.g., selective laser melting, electron beam melting), directed metal deposition methods (e.g., powder deposition and wire-based methods).

Background Art

[0002] Currently, there is a tendency to transfer nickel-based superalloys successfully manufactured in cast or forged form to the AM process. However, this has proven to be almost inappropriate. This is because many of the material properties required to facilitate processing in the AM process are not met by the above alloys, making processing significantly more difficult and resulting in materials without the expected structural integrity.

[0003] In particular, developing alloys with a high γ´ volume fraction for additive manufacturing processes has been a major challenge. This is because these alloys are often classified as "non-weldable". Usually, these alloys are processed by investment casting. Examples of common alloys used in the investment casting process are listed in Table 1. Table 1 shows the nominal composition in mass % of conventional high γ´ volume fraction alloys.

[0004] Research on applying the alloys listed in Table 1 to the additive manufacturing method has been widely conducted. Although the alloys may be "non-weldable" and difficult to process, it has been shown that specific defect mechanisms can be limited. For example, these alloys are susceptible to strain age cracking, but careful control of AM conditions (scan plan, heat input, etc.) and post-treatment conditions (heat treatment window and heating rate) during processing can minimize the risk of strain age cracking to a certain extent. Similarly, the strain age cracking resistance can be improved by designing the shape of the part to reduce the influence of stress concentration features such as notches.

[0005] Strain age cracking is caused by two main factors: residual strain and γ' precipitation. The magnitude of the residual strain can be easily reduced by preheating the powder bed, which minimizes the temperature range of thermal contraction. Illston (U.S. Patent No. 9,352,421 B2) shows that process control (especially the use of thin powder layers and intentionally overlapped laser scans) can minimize the accumulation of residual strain and thus improve the printability of high γ' superalloys. Also, Etter et al (U.S. Patent No. 9,670,572 B2) shows that the risk of strain age cracking can be reduced by heating up to the stress relief heat treatment temperature after AM particularly rapidly.

[0006] Another mechanism that cannot be easily reduced by process optimization is hot cracking. Hot cracking occurs in the final stage of the solidification process and strongly depends on the chemical properties of the alloy. SUMMARY OF THE INVENTION PROBLEM TO BE SOLVED BY THE INVENTION

[0007] An object of the present invention is to develop an alloy with a high γ' volume fraction that is particularly resistant to the mechanism of hot cracking by adjusting the chemical properties of the alloy so as to overcome the undesirable damage mechanism of hot cracking.

[0008] In combination with this improved workability, the alloys of the present invention preferably have a very high level of oxidation resistance (achieved by having a level of aluminum sufficient to form a protective aluminum-based oxide scale) and a high level of creep resistance.

[0009] TABLE 1 MEANS FOR SOLVING THE PROBLEM

[0010] According to the present invention, aluminum of 4.0 to 5.6% by mass, titanium of 0.0 to 1.0% by mass, niobium of 0.0 to 4.0% by mass, tantalum of 0.0 to 11.9% by mass, tungsten of 2.0 to 12.7% by mass, molybdenum of 0.0 to 3.0% by mass, cobalt of 0.0 to 22.0% by mass, chromium of 6.0 to 16.7% by mass, carbon of 0.02 to 0.35% by mass, boron of 0.001 to 0.2% by mass, zirconium of 0.00 to 0.5% by mass, rhenium of 0.0 to 3.0% by mass, ruthenium of 0.0 to 2.0% by mass, iridium of 0.0 to 3.0% by mass, vanadium of 0.0 to 0.5% by mass, palladium of 0.0 to 1.0% by mass, platinum of 0.0 to 1.0% by mass, silicon of 0.0 to 0.5% by mass, yttrium of 0.0 to 0.1% by mass, lanthanum of 0.0 to 0.1% by mass, cerium of 0.0 to 0.1% by mass, sulfur of 0.0 to 0.003% by mass, manganese of 0.0 to 0.25% by mass, magnesium of 0.0 to 0.1% by mass, iron of 0.0 to 4.0% by mass, copper of 0.0 to 0.5% by mass, and hafnium of 0.0 to 1.0% by mass are included, with the balance being nickel and unavoidable impurities, and the mass percentages of niobium and tantalum contained in the alloy being W Nb and W Ta respectively. Then, a nickel-based alloy composition satisfying the following formula is provided. 1.1 ≦ 0.3W Nb + 0.15W Ta Such a nickel-based alloy has excellent hot cracking resistance, along with appropriate creep resistance, appropriate strength, appropriate stress relaxation cracking resistance, appropriate microstructure stability, an acceptable solidification temperature range, an acceptable density, and an acceptable cost.

[0011] In one embodiment, when the mass percentages of niobium and tantalum contained in the alloy are W Nb and W Ta respectively, the following formula is satisfied. 1.15 ≦ 0.3W Nb + 0.15W Ta Preferably, the following formula is satisfied. 1.2 ≦ 0.3W Nb + 0.15W Ta More preferably, the following formula is satisfied. 1.4 ≤ 0.3W Nb + 0.15W Ta Even more preferably, the following formula is satisfied. 1.8 ≤ 0.3W Nb + 0.15W Ta In such an alloy, the hot cracking resistance is further improved.

[0012] In one embodiment, when the mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy are W Al , W Ti , W Nb and W Ta respectively, the following formula is satisfied. 5.6 ≤ W Al + 0.5W Ti + 0.3W Nb + 0.15W Ta ≤ 7.0 Preferably, the following formula is satisfied. 5.6 ≤ W Al + 0.5W Ti + 0.3W Nb + 0.15W Ta ≤ 6.5 In such an alloy, the stress relaxation cracking resistance is improved.

[0013] In one embodiment, when the mass percentages of tungsten and molybdenum contained in the alloy are W W , W Mo respectively, the following formula is satisfied. W W + 0.65W Mo ≥ 4.0 Preferably, the following formula is satisfied. W W + 0.65W Mo ≥ 6.0 More preferably, the following formula is satisfied. W W + 0.65W Mo ≥ 8.0 In such an alloy, the creep resistance is improved.

[0014] In one embodiment, chromium included in the nickel-based alloy composition is 8.0% or more, preferably 10.0% or more by mass%. In such an alloy, oxidation resistance and corrosion resistance are improved.

[0015] In one embodiment, chromium included in the nickel-based alloy composition is 14.7% or less, preferably 13.8% or less by mass%. In such an alloy, fine structure stability is improved.

[0016] In one embodiment, tantalum included in the nickel-based alloy composition is 9.8% or less, preferably 9.2% or less, more preferably 7.1% or less by mass%. Such an alloy has a low density and optionally the tungsten content can be increased. The increase in the tungsten content helps to improve creep resistance.

[0017] In one embodiment, molybdenum included in the nickel-based alloy composition is 2.0% or less, preferably 1.8% or less by mass%. In such an alloy, hot cracking resistance is improved.

[0018] In one embodiment, titanium included in the nickel-based alloy composition is 0.5% or less, preferably 0.1% or less by mass%. Such an alloy has better oxidation resistance.

[0019] In one embodiment, tungsten included in the nickel-based alloy composition is 10.7% or less, preferably 8.7% or less by mass%. In such an alloy, fine structure stability is improved.

[0020] In one embodiment, niobium included in the nickel-based alloy composition is 3.0% or less by mass%. In such an alloy, oxidation resistance is improved.

[0021] In one embodiment, one or both of platinum and palladium included in the nickel-based alloy composition is 0.5% or less by mass%. In such an alloy, the cost is further reduced.

[0022] In one embodiment, the aluminum included in the nickel-based alloy composition is 4.5% or more by mass%. In such an alloy, the corrosion resistance is improved.

[0023] In one embodiment, the aluminum included in the nickel-based alloy composition is 5.3% or less, preferably 5.1% or less, more preferably 4.6% or less by mass%. In such an alloy, in combination with the improvement of hot cracking resistance, strain age cracking and toughness are improved.

[0024] In one embodiment, the cobalt included in the nickel-based alloy composition is 15.0% or less by mass%. In such an alloy, the solidification temperature range is lowered.

[0025] In one embodiment, the tantalum included in the nickel-based alloy composition is 1.1% or more, preferably 3.7% or more, more preferably 5.8% or more by mass%. In such an alloy, the amount of niobium can be reduced without reducing the hot cracking resistance, or the hot cracking resistance can be further improved by combining with a higher level of niobium.

[0026] In one embodiment, the tungsten included in the nickel-based alloy composition is 2.7% or more, preferably 2.8% or more, more preferably 4.7% or more, even more preferably 5.1% or more, most preferably 5.2% or more by mass%. In such an alloy, the creep resistance is improved.

[0027] In one embodiment, let the mass% of tantalum and tungsten contained in the alloy be W Ta and W W respectively, then the following formula is satisfied W W +W Ta ≦13.9 Preferably, the following formula is satisfied. W W +W Ta ≦11.8 In these alloys, the density is lowered.

[0028] In one embodiment, when the mass percentages of chromium, molybdenum, and tungsten contained in the alloy are W Cr , W Mo , and W W , respectively, the following equation is satisfied. W W + W Cr + 1.7W Mo ≤ 18.7 Preferably, the following equation is satisfied. W W + W Cr + 1.7W Mo ≤ 17.8 In such an alloy, the stability of the microstructure is improved.

Brief Description of the Drawings

[0029]

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Embodiments for Carrying Out the Invention

[0030] Conventionally, nickel-based superalloys have been designed based on empiricism. Therefore, the chemical composition of nickel-based superalloys has been identified through time-consuming and expensive experimental development, including small-scale processing of limited amounts of materials and subsequent characterization of their behavior. Subsequently, an alloy composition that has been found to exhibit the best, i.e., the most desirable, combination of properties is adopted. The fact that there are a number of alloy element groups capable of achieving this combination indicates that these alloys are not fully optimized and that there is a high likelihood that more improved alloys exist.

[0031] In superalloys, generally, chromium (Cr) and aluminum (Al) are added to impart oxidation / corrosion resistance, and cobalt (Co) is added to improve resistance to sulfidation. For creep resistance, molybdenum (Mo), tungsten (W), and cobalt are introduced because these elements inhibit the thermally activated processes (e.g., dislocation climb) that determine the rate of creep deformation. To increase static and fatigue strength, aluminum (Al), tantalum (Ta), niobium (Nb), and titanium (Ti) are introduced because these elements promote the formation of the precipitation hardening phase gamma prime (γ´). This precipitate is coherent with the face-centered cubic (FCC) matrix phase called gamma (γ).

[0032] In this specification, a model-based approach used to identify new grades of nickel-base superalloys is described in terms of the term "alloy design" (ABD) method. This approach utilizes a framework of computational materials models to estimate design-related properties over a very wide composition range. In principle, this alloy design tool enables the solution of so-called inverse problems. That is, the optimal alloy composition that most satisfies the specified design constraints can be identified.

[0033] The first step in the design process is to define an element table and the upper and lower limits of the composition constraints associated with that element table. In the present invention, the composition limits for each element when adding each element, called the "alloy design region", are considered. These composition limits are detailed in Table 2. Table 2 shows the alloy design region in mass % investigated using the "alloy design" method.

[0034]

Table 2

[0035] The balance is nickel. The levels of carbon, boron, and zirconium were fixed at 0.06%, 0.015%, and 0.06%, respectively.

[0036] The second step is performed based on thermodynamic calculations to calculate the phase diagram and thermodynamic properties of a specific alloy composition. This is often called the CALPHAD method (CALculation of PHAse Diagrams). By performing these calculations at the typical operating temperature (900 °C) of the new alloy, information about the phase equilibrium (microstructure) can be obtained.

[0037] The third stage involves identifying an alloy composition with the desired microstructure. In the case of nickel-based superalloys that require excellent resistance to creep deformation, the creep rupture life is gradually improved as the volume fraction of the precipitation hardening phase γ´ increases. The range of the volume fraction of γ´ for which the creep rupture life is most beneficial is 60 - 70%. When the volume fraction of γ´ exceeds 70%, a decrease in creep resistance is observed.

[0038] Also, the γ / γ´ lattice misfit needs to follow a small value, either positive or negative, because it loses coherency. Therefore, the limit depends on the absolute value of that value. The lattice misfit δ is defined as the mismatch between the γ phase and the γ´ phase and is obtained by the following equation.

[0039]

Equation

[0040] Here, α γ and α γ´ are the lattice constants of the γ phase and the γ´ phase.

[0041] Alloys based on an inappropriate microstructure are also rejected by the estimated susceptibility to the topologically close-packed (TCP) phase. By using CALPHAD modeling in this calculation, the formation of harmful TCP phases sigma (σ), rho (Ρ), and mu (μ) is predicted.

[0042] Therefore, according to this model, all the compositions within the design space for which the calculated result of the volume fraction of γ´ becomes the desired value are identified. In these compositions, the lattice distortion of γ´ is less than a predetermined absolute value, and the total volume fraction of the TCP phase is less than a predetermined magnitude.

[0043] In the fourth step, for the identified alloy compositions remaining in the dataset, a merit index is estimated. Examples of the merit index include a creep merit index (indicating the creep resistance of an alloy based on the average composition only), a strength merit index (indicating an alloy’s precipitation yield strength based on the average composition only), a solution merit index (indicating the solution yield strength of an alloy based on the average composition only), density, and cost.

[0044] In the fifth step, the calculated merit index is compared with the constraints for the desired behavior, and these design constraints are regarded as the boundary conditions for the problem. All compositions that do not satisfy the boundary conditions are rejected. At this stage, the size of the test dataset becomes very small.

[0045] The final sixth step involves analyzing the remaining dataset of compositions. This analysis can be performed in various ways. One way is to classify, via the database, the alloy for which the merit index shows the maximum value. The alloy for which the merit index shows the maximum value is, for example, the lightest alloy, the most creep-resistant alloy, the most oxidation-resistant alloy, and the least expensive alloy. Or, alternatively, the database can be used to determine the relative performance trade-offs resulting from different combinations of properties.

[0046] Seven examples of the merit index are described.

[0047] The first merit index is the creep merit index. The most important observation is that the time-dependent deformation (i.e., creep) of nickel-based superalloys occurs by dislocation creep associated with limited initial activity in the γ phase. Therefore, as the proportion of the γ' phase increases, dislocation segments are rapidly fixed at the γ / γ' interface. The rate-determining step is the detachment of the trapped configuration of dislocations from the γ / γ' interface. It depends on the local chemistry (in this case, the composition of the γ phase) that has a significant impact on the creep properties with respect to the alloy composition.

[0048] The microstructure model based on physics is invoked for the accumulation rate of creep strain ε · when the load is uniaxial and along the <001> crystallographic direction. The set of equations is as follows.

[0049]

Number

[0050] Here, ρ m is the mobile dislocation density, φ p is the volume fraction of the γ' phase, and ω is the width of the matrix channel. The terms σ and Τ are the applied stress and temperature, respectively. The terms b and k are the Burgers vector and the Boltzmann constant, respectively. The term K CF is the constraint coefficient.

[0051]

Number

[0052] The term K CF indicates the proximity of cubic particles in these alloys. Equation 3 shows a dislocation multiplication process that requires the estimation of the multiplication parameter C and the initial dislocation density. The term D eff is the effective diffusion rate that controls the climb process at the particle / matrix interface.

[0053] Note that in the above content, the composition dependence is related to the two terms φ p and D effresults from. Therefore, assuming that the microstructure is constant (most of the microstructure is controlled by heat treatment), φ p is fixed, and the dependence on chemical composition is caused by D eff . For the purposes of the alloy design modeling described herein, it can be seen that it is not necessary to perform the complete integration of Equations 2 and 3 for each prototype alloy composition. Instead, the primary merit index M creep that needs to be maximized is used. M creep is obtained by the following equation.

[0054]

Equation

[0055] where x i is the atomic fraction of solute i in the γ phase. D i ~ is the appropriate interdiffusion coefficient.

[0056] The second merit index is the strength merit index. In the case of high-nickel-based superalloys, most of the strength is derived from the precipitate phase. Therefore, optimizing the alloy composition to maximize the precipitation strength is an important design consideration. Based on the hardening theory, a strength merit index M strength is proposed. This index takes into account the maximum possible precipitation strength (determined as the point where dislocation shear transfer occurs from weak bonds to strong bonds) and is approximated using the following equation.

[0057]

Equation

[0058] where M- is the Taylor coefficient, γ APB is the antiphase boundary (APB) energy, φ p is the volume fraction of the γ´ phase, and b is the Burgers vector.

[0059] From Equation (5), it is clear that the defect energy in the γ´ phase (e.g., the antiphase boundary APB energy) has a significant impact on the deformation behavior of nickel-based superalloys. It has been found that increasing the APB energy improves the mechanical properties including tensile strength and resistance to creep deformation. Research on the APB energy has been carried out for many Ni-Al-X systems using density functional theory. From this research, the influence of ternary elements on the APB energy of the γ´ phase was calculated, and a linear superposition of the effects of adding each ternary element was assumed when considering the complex multi-component system. As a result, the following equation was derived.

[0060] [Number]

[0061] Here, x Cr , x Mo , x W , x Ta , x Nb and x Ti represent the atomic % concentrations of chromium, molybdenum, tungsten, tantalum, niobium, and titanium in the γ´ phase, respectively. The composition in the γ´ phase is determined by phase equilibrium calculations.

[0062] The third merit index is density. The density ρ was calculated using the simple rule of mixtures and a correction factor. Here, ρ i is the density of a given element, and x i is the atomic fraction of the alloying element.

[0063] [Number]

[0064] The fourth merit index is cost. To estimate the cost of each alloy, the simple rule of mixtures was applied. Here, the cost of each alloy was calculated using the mass fraction x i of the alloying element multiplied by the current (2016) raw material cost c i of the alloying element.

[0065]

Number

[0066] This estimation assumes that the processing cost is the same for all alloys. That is, the product yield is not affected by the composition.

[0067] The fifth merit index is based on the exclusion of alloy candidates based on the inappropriate microstructure based on the susceptibility to the TCP phase. To do this, the d-orbital energy level of the alloying element (referred to as Md) is used to determine the total effective Md level according to the following formula.

[0068]

Number

[0069] Here, x i represents the mole fraction of element i contained in the alloy. The higher the value of Md, the higher the possibility of TCP formation.

[0070] The sixth merit index is the strain aging crack index. The performance of alloys produced by additive manufacturing is related to their scientific composition. This index has been developed based on empirical observations associating alloy composition with the weldability of nickel-based superalloys from the perspective of susceptibility to strain aging cracking. In this relationship, since the density of titanium is approximately twice that of aluminum, multiplying the titanium content by a factor of 0.5 converts it to an "equivalent aluminum amount". In fact, the additive manufacturing process of metal alloys is a continuous welding process. Previous observations that associated weldability only with the aluminum and titanium contents have been applied. A modification is included to account for the influence of tantalum and niobium, which behave similarly to aluminum and titanium during alloy aging. Similar to titanium, a factor is assigned to convert the addition amounts of these elements to an "equivalent aluminum amount". That is, the correction factors for niobium and tantalum (determined from their densities relative to aluminum) are 0.3 and 0.15, respectively. The strain aging index is given by the following equation.

[0071]

Equation

[0072] where W Al 、W Ti 、W Nb and W Ta represent the mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy, respectively. The lower the value of this strain aging index, the better the response to the additive manufacturing process.

[0073] The seventh merit index is based on the solidification behavior of alloy candidates predicted by the Scheil-Gulliver model. This is for ranking the susceptibility to hot cracking based on composition. In this approach, the temperature range at the final stage of solidification (solid fraction of 90 - 99%) is regarded as representing the region where the alloy is vulnerable to hot cracking. This is because at this stage, the supply of liquid is likely to be restricted by the bridged network of solid materials. The temperature range with a solid fraction of 40 - 90% is regarded as a safe region. This is because at this stage, the restriction on the supply of liquid is significantly relaxed. The temperature range with a solid fraction of less than 40% is considered inappropriate because the liquid is dominant.

[0074] According to the ranking system used for alloy casting by Clyne and Davis, the hot cracking index is defined as follows, as the ratio of the vulnerable temperature range to the safe temperature range.

Equation

[0075] The lower the value of this index, the lower the risk of hot cracking.

[0076] Using the ABD method described above, the alloy composition of the present invention was identified. The design intention of this alloy is to develop a superalloy that has high creep resistance (achieved by increasing the γ' volume fraction) and improved workability by additive manufacturing (achieved by improving the resistance to strain-age cracking and the resistance to the mechanism of hot cracking) compared to other alloys with a high γ' volume. In addition to these attributes, by having a sufficient level of aluminum to form a protective aluminum-based oxide scale, good oxidation resistance is achieved, and other important material properties including microstructural stability and alloy density are optimized.

[0077] The material properties of the conventional compositions (listed in Table 1) (determined using the ABD method) are listed in Table 3. In association with the predicted properties listed for these alloys, the design of new alloys was considered. Table 3 shows the calculated phase fractions and merit indices created by the "alloy design" software. This is the result for the nickel-based superalloys listed in Table 1.

[0078] The design principle of the new alloy will be described below.

[0079]

Table 3

[0080] Alloys with a high strain age cracking index are known to be difficult to process based on welding literature, and the same tendency is generally observed in AM. Reducing the strain age merit index reduces the tendency of this defect mechanism, but the high-temperature strength (based on the viewpoints of creep strength and tensile strength) and the oxidation resistance (based on the viewpoint of the formed protective oxide scale) are dependent on a relatively high content of γ'-forming elements that make it difficult to suppress the improvement of the strain age index. To overcome this, strain age cracking can be managed by other methods. For example, by careful control of AM conditions (scan plan, heat input, etc.) during processing and post-treatment conditions (heat treatment window and heating rate), the risk related to strain age cracking can be limited. Similarly, by designing the shape of the part to reduce the stress concentration function that causes local strain age cracking, the risk of strain age cracking can be reduced. See, for example, US9352421B2 and US9670572B2.

[0081] Unlike strain aging cracking, another manufacturing defect that can occur during the additive manufacturing process is hot cracking. Hot cracking can occur during the solidification process. The mechanism of this manufacturing defect is caused by the change in the liquid composition during the solidification of the alloy. The risk of hot cracking occurring cannot be managed in the same way as strain aging cracking. Carefully controlling the AM conditions during processing may help to some extent, but for example, the use of post-processing or design changes does not affect this mechanism. The most direct way to improve resistance to hot cracking is to change the chemical properties of the alloy. Improving resistance to hot cracking significantly improves the workability of the alloy for additive manufacturing, even when it is necessary to allow a high strain aging cracking index. The object of the present invention is to identify an alloy (see FIG. 1) that has a low hot cracking index value compared to other alloys and a high strain aging index within the range where strain aging cracking can be managed by processing conditions.

[0082] FIG. 2 shows the relationship between the addition amounts of aluminum element, niobium element, and tantalum element, which are predominantly added to form the γ' phase and control the γ' volume fraction.

[0083] Elements that form the γ' phase may reduce the ease of processing the alloy by AM because the strain aging merit index increases (Equation 10). Therefore, it is necessary to optimize the combination of these elements to provide a desirable balance between limiting the possibility of strain aging cracking during AM processing and the γ' volume fraction (providing strength from the viewpoints of creep resistance and tensile strength).

[0084] To achieve a favorable strain aging cracking index, the contents of aluminum, titanium, niobium, and tantalum need to satisfy the following constraints. f(SAC)=W Al +0.5W Ti +0.3W Nb +0.15W Ta

[0085] Here, f(SAC) is a numerical value of 7.0 or less, and W Al, W Ti , W Nb and W Ta respectively indicate the mass percentages of aluminum, titanium, niobium, and tantalum contained in the alloy. Numerical values of 7.0 or less are selected to be equal to or better than alloy IN713C (see Table 3). Setting the value of f(SAC) above 7 is not preferable because alloy processing is restricted from the perspective of stress-rupture cracking. To facilitate processing by AM, it is desirable to lower the stress-rupture merit index. That is, W Al + 0.5W Ti + 0.3W Nb + 0.15W Ta ≤ 6.5 is preferable.

[0086] The intended application temperature of this new alloy is up to 1000 °C in high-oxidizing and corrosive environments. Examples of application fields include the high-temperature section of a gas turbine engine or within the exhaust system of an internal combustion engine.

[0087] To achieve desirable oxidation performance, it is desirable that the alloy of the present invention forms a protective aluminum-based oxide scale (Al 2 O 3 ). Such alloys are stable above 1000 °C, in contrast to those based on Cr 2 O 3 . Alloys such as IN738 and IN792 have a relatively high γ´ volume fraction and extremely excellent creep resistance, but due to their relatively low aluminum contents (3.4 mass% and 3.2 mass% respectively), they cannot form an aluminum-based protective oxide scale. Instead, a less protective chromium-based oxide scale is formed. The aluminum-based oxide scale is excellent in adhesion and thermal stability, while in the chromium-based protective oxide scale, chromium volatilization can occur near 1000 °C, resulting in a less protective oxide. To produce an aluminum-based protective oxide scale, an aluminum content of at least 4.0 mass%, more preferably 4.5 mass% or more, is required. This further improves the formation of the alumina scale and further enhances the oxidation resistance.

[0088] Alloying with titanium is known to reduce oxidation performance due to the formation of titanium oxide. This is known to be significantly detrimental to oxidation performance. In the case of the alloys of the present invention, the amount of titanium used is limited to up to 1.0% by mass, preferably up to 0.5% by mass, and most preferably up to 0.1% by mass due to the adverse effect on high-temperature oxidation resistance. Also, since titanium has a relatively high coefficient with respect to strain aging cracking (Equation 10), it is beneficial to limit titanium to reduce the tendency of strain aging cracking. On the other hand, by setting the minimum amount of titanium to 0.1% by mass, the density can be limited and the formation of carbides can be advantageously promoted.

[0089] Alloying with niobium is also known to have an adverse effect on oxidation. This is because niobium forms grain boundary carbides. These grain boundary carbides are particularly harmful to oxidation-assisted cracking mechanisms where damage can accumulate along grain boundaries, such as under low-cycle fatigue conditions, creep fatigue conditions, or during high-temperature creep. However, since the addition of niobium is beneficial for improving the resistance to high-temperature cracking (to be described later with reference to FIG. 3), its use is limited to up to 4.0% by mass, more preferably limited to 3.0% by mass or less.

[0090] By adding chromium, the formation of a protective alumina oxide scale is promoted. In particular, chromium is desirable for improving resistance to high-temperature corrosion. Alloys such as IN738 and IN792 contain relatively high levels of chromium (16.0 wt% and 12.7 wt% respectively). This is mainly for resistance to high-temperature corrosion, but since they do not form a protective aluminum-based oxide, the oxidation rate at very high temperatures is slow, and therefore the maximum operating temperature is limited. These alloys also have a lower resistance to creep compared to IN713C (Figure 4), so the upper limit of the operating temperature is restricted. The alloy of the present invention needs to contain 6.0 wt% or more of chromium. In order to achieve good high-temperature corrosion resistance, a chromium level of 6.0 wt% or more is desirable. More preferably, the chromium content is 8.0 wt% or more in order to provide high-temperature corrosion resistance equivalent to CM247LC. Even more preferably, chromium is present in an amount of 10.0% or more. This further enhances the corrosion resistance compared to the alloy CM247LC.

[0091] Molybdenum is known to significantly reduce the high-temperature corrosion resistance of nickel superalloys. Such molybdenum is a selective additive, but preferably by setting a minimum value of 0.1 wt% or more, and further 0.5 wt% or more, it helps to reduce the high-temperature corrosion resistance. In order to achieve good corrosion resistance, it is desirable to limit molybdenum to 3.0 wt% or less. More preferably, molybdenum is limited to 2.0 wt% or less. Since alloys IN738 and IN792 are known to have very excellent corrosion resistance, most preferably molybdenum is limited to 1.8 wt% or less.

[0092] Based on the minimum aluminum content (4.0 wt%) and the desirability of f(SAC) < 7, and the desirability of limiting the titanium content of the alloy to 0.1 wt% or less (for example, 0.0 wt%), 0.3W Nb +0.15W TaIt is necessary to limit the total of tantalum element and niobium element according to the relationship to 3.0 or less. Therefore, it is necessary to set the upper limits of niobium and tantalum to 9.7% by mass and 20.0% by mass, respectively. Preferably, in order to improve the balance between oxidation resistance (aluminum 4.5% by mass or more) and stress-relaxation cracking resistance (f(SAC) < 6.5), 0.3W Nb +0.15W Ta is limited to 2.5 or less. Therefore, it is more preferable that the upper limits of niobium and tantalum are 8.1% by mass and 16.6% by mass, respectively. When aluminum is 5.0% by mass and f(SAC) < 6.5, the balance between resistance to stress-relaxation cracking and oxidation resistance is most preferable. That is, 0.3W Nb +0.15W Ta is most preferably limited to 2.0 or less. Therefore, the upper limits of niobium and tantalum need to be 6.5% by mass and 13.3% by mass, respectively. However, as will be described later, the alloys of the present invention actually require even lower levels of niobium and tantalum. At the most preferable level of tantalum (7.1% by mass or less), niobium is an essential element to achieve the required hot cracking index. In order to improve hot cracking resistance, niobium of 0.05% by mass or more, 0.5% by mass or more, and further 1.0% by mass or more is desirable. All of the following examples include at least such a level of niobium.

[0093] In FIG. 2, dotted lines indicating various limits of stress-relaxation cracking are shown. It can be seen that in order to produce an alloy with a stress-relaxation index of 7 or less, it is preferable to limit the γ' volume fraction at an equilibrium temperature of 900 °C to 0.63. Preferably, the γ' volume fraction is based on the more preferable value of stress-relaxation index W Al +0.5W Ti +0.3W Nb +0.15W Ta ≦6.5 and is limited to 0.56.

[0094] The desirable minimum requirement for the γ´ volume fraction is 0.42. This will be described later in connection with achieving the necessary creep resistance using FIGS. 4 and 5. To achieve the desirable volume fraction of γ´, the contents of aluminum, titanium, niobium, and tantalum need to satisfy the following constraints. f(γ´)=W Al +0.5W Ti +0.3W Nb +0.15W Ta

[0095] Here, f(γ´) is a numerical value in the range of 5.6 to 7.0, and at this time, an alloy with a desirable γ´ fraction of 0.42 to 0.63 is produced. When f(γ´) is a numerical value in the range of 5.6 to 6.5, an alloy with a γ´ ratio of 0.42 to 0.56 is produced, and an alloy having an improved combination of high creep strength and high AM workability is obtained. Although the improvement of the creep resistance of the alloy will be described later, the γ´ ratio is preferably 0.43 or more, more preferably 0.45 or more, and most preferably 0.50 or more. As a result, the desirable numerical values of f(γ´) are 5.7 or more, 5.8 or more, and 6.1 or more, respectively.

[0096] Platinum elements and palladium elements exhibit the same behavior as tantalum, titanium, and niobium. That is, these are γ´-forming elements that increase the antiphase boundary energy. These elements can be selectively added to the alloy instead of the tantalum element, titanium element, and niobium element. The advantages of this may include an improvement in high-temperature corrosion resistance. The "aluminum equivalent amounts" of platinum and palladium require correction coefficients of 0.125 and 0.225, respectively (determined from the density relative to aluminum). However, since the addition of these elements is costly, the addition amounts of these elements can be limited. Therefore, these elements can each be present at a maximum of 1.0 mass%, preferably limited to 0.5 mass% or less, and most preferably limited to 0.1 mass% or less. This range provides an optimal balance between cost and improvement in corrosion resistance. To perform good processing by additive manufacturing, it is preferable to satisfy the following formula. W Al+0.5W Ti +0.3W Nb +0.15W Ta +0.125W Pt +0.225W Pd ≦7.0 Preferably, it satisfies the following formula. W Al +0.5W Ti +0.3W Nb +0.15W Ta +0.125W Pt +0.225W Pd ≦6.5

[0097] Here, W Pt and W Pd respectively represent the mass percentages of platinum and palladium contained in the alloy.

[0098] The tendency of the alloy to form hot cracks is determined from the perspective of the hot crack index (Equation 11). This hot crack mechanism is mainly caused by elements that segregate to the liquid phase during solidification. In the investigated design region (Table 2), niobium and tantalum are the elements that segregate most strongly to the liquid phase, so they have the strongest influence on hot cracking.

[0099] Figure 3 shows the hot crack index as a function of aluminum, niobium, and tantalum elements. These elements promote strain age cracking (as explained in relation to Figure 1, a balance between strain age cracking and hot cracking is necessary for the best AM processability). It can be seen that niobium and tantalum strongly affect the hot crack index. The alloys listed in Table 1 tend to develop hot cracks during AM, and alloy IN738 has the least tendency to crack with an index of 1.8. The hot crack indices of alloys that are very well processed by AM and do not show hot cracks, such as alloy 718 and alloy 625, are 1.0 or less (Figure 1). For alloys with significantly excellent resistance to hot cracking, an index of 1.5 or less is useful, and from Figure 3, it is judged that the niobium and tantalum contents need to satisfy the following constraints to achieve the desired hot crack index. f(HCI)=0.3WNb +0.15 W Ta

[0100] Here, f(HCI) is a numerical value, and by setting it to a value of 1.1 or more (or 1.10 or more), a hot cracking index of 1.5 or less is achieved. On FIG. 2, a line indicating a hot cracking index of 1.5 is superimposed. In combination with a strain aging cracking index of 7.0 or less, it can be seen that aluminum must be limited to 5.6 mass% or less in order to achieve a hot cracking index of 1.5 or less. More preferably, since it is desirable to set the hot cracking index to 1.5 or less in combination with a strain aging cracking index of 6.5 or less, it is preferable to limit aluminum to 5.3 mass%. It is more preferable to set the maximum level of aluminum to 5.3 mass% or less because the risk of strain aging cracking is reduced and all alloy examples described later can fall within this range. In order to improve the alloy's resistance to oxidation-assisted cracking, it is preferable to limit niobium to 3.0 mass% or less. When niobium is set to a preferable level of 3.0 mass% in order to achieve a hot cracking index of 1.5 or less in combination with a strain aging index of 6.5 or less, the alloy preferably contains at least 1.1 mass% of tantalum in order to achieve the desired value of the hot cracking index.

[0101] In order to enhance the resistance to hot cracking, it is more preferable to set the hot cracking index to 1.0 or less. That is, f(HCI) preferably becomes 1.8 or more. It is also desirable for f(HCI) to be a value of 1.15 or more or 1.2 or more or 1.4 or more. On FIG. 2, lines indicating hot cracking indices of 1.0 and 1.5 are superimposed (HCI = 1 and HCI = 1.5). In combination with a strain aging cracking index of 7.0 or less, aluminum is limited to 5.1 mass% or less in order to achieve a hot cracking index of 1.0 or less. More preferably, since it is desirable to set the hot cracking index to 1.0 or less in combination with a strain aging cracking index of 6.5 or less, it is preferable to limit aluminum to 4.6 mass% or less.

[0102] Based on the maximum niobium content (4.0% by mass) to achieve a preferred value of f(HCI) (1.8 or more), the tantalum content is preferably at least 3.7% by mass, more preferably, the niobium is limited to 3.0% by mass or less, and thus the tantalum content is preferably 5.8% by mass or more.

[0103] The addition of cobalt has the effect of lowering the γ' solvus temperature. A decrease in the γ' solvus temperature is desirable because it lowers the temperature at which γ' precipitation occurs. This is advantageous for reducing the rate of strain age hardening because the rate of strain age hardening depends on γ' precipitation. Lowering the γ' solvus also improves the ability to perform a solution treatment. The ability to perform a solution treatment is necessary to homogenize the distribution of elemental species that occurs after the AM process, to change certain microstructural features such as increasing the particle size, and to dissolve coarse γ' precipitates that do not provide a significant strengthening effect. By rapidly cooling from the solution treatment temperature, a fine dispersion of γ' particles can be achieved, which helps to improve the mechanical properties. Therefore, it is desirable to set the minimum level of cobalt at 8.0% by mass or more, and all alloy examples described below have at least such a level of cobalt. A more preferred minimum level of cobalt is 9.0% by mass, and an even more preferred limit is 10.0% by mass or more.

[0104] However, as the cobalt content increases, the shell solidification temperature range of the alloy increases (Figure 13). An increase in the solidification range is associated with an increase in the time to solidification, and there may be a risk of solidification cracking when the alloy is in a semi-solid state, so it is desirable to limit the solidification range. It is desirable to set the target freezing range to 300°C or less, that is, cobalt up to 22.0% by mass is acceptable. To match the freezing range of IN792, it is desirable to set the cobalt content to 15.0% by mass or less.

[0105] The relationship between the γ´ volume fraction and the creep merit index related to creep resistance (from the perspective of temperature capability at 137 MPa, using contours normalized to IN713) is shown in Fig. 4. Increasing both parameters increases the creep resistance, and the sensitivity to each parameter is determined. The positions of the alloys listed in Table 3 are shown in Fig. 4. The object of the present invention is to have creep performance equivalent to IN713C, and more preferably, an improvement of 25 °C is desirable.

[0106] For the alloy of the present invention, it is desirable that the creep merit index is higher than that of alloy IN713C. Therefore, in the case of the alloy of the present invention, it is desirable to set the creep merit index to 6.0×10 -15 m -2 s. The addition amounts of the elements necessary to achieve this level of creep merit index will be confirmed in the following section with reference to Fig. 5. The maximum achievable creep merit index is determined by the fact that in order to maintain a stable microstructure essentially free of the TCP phase, the alloy needs to contain at least 6.0 mass% chromium for corrosion resistance (see Fig. 5). To obtain creep resistance equivalent to IN713C, it is desirable to set the γ´ volume fraction to 0.42. More preferably, an improvement of at least 25 °C in the creep resistance of IN713C is desirable. Therefore, especially when the creep merit index is 6.0×10 -15 m -2 s, the γ´ volume fraction is preferably 0.50, which corresponds to 6.1 or more in the numerical value of f(γ´).

[0107] Slow-diffusion elements distributed in the gamma matrix phase have the strongest influence on the creep merit index. This creep merit index is calculated based on the composition of the gamma phase at the equilibrium temperature of 900 °C. Tungsten is the element with the slowest diffusion in the alloy design region listed in Table 2, followed by molybdenum. The influence of the elements tungsten and molybdenum on creep resistance is shown in Fig. 5. From Fig. 5, it is judged that the change in the creep merit index is related to the total of the molybdenum element and the tungsten element based on the following formula. f(CMI)=W W+0.65 W Mo

[0108] Here, f(CMI) is a numerical value, and Ww and W Mo respectively represent the mass % of tungsten and molybdenum in the alloy. To achieve a desirable creep merit index of 6.0×10 -15 m -2 s, a value of f(CMI) of 4.0 or more is desirable. Based on the upper limit of molybdenum (3.0 mass %), the tungsten content needs to be at least 2.0 mass %. More preferably, since molybdenum is limited to 2 mass %, it is preferable to set the tungsten content to 2.7 mass % or more. Most preferably, since molybdenum is limited to 1.8 mass %, it is most preferable to contain tungsten at least 2.8 mass % or more. A more desirable level of f(CMI) is 6.0 or more, preferably 8.0 or more.

[0109] For an even better combination of creep resistance and stress rupture crack resistance, it is desirable to improve the temperature capacity by 25 °C compared to IN713C in combination with a stress rupture crack index of 6.5 or less. As a result, the maximum value of the γ' volume fraction is limited to 0.56, and thus a creep merit index of 6.90×10 -15 m -2 s or more is required. To achieve a creep merit index of 6.90×10 -15 m -2 s or more, the value of f(CMI) needs to be 6.4 or more. Based on the upper limit of molybdenum (3 mass %), the tungsten content needs to be at least 4.7 mass %. More preferably, molybdenum is limited to 2.0 mass %, that is, it is preferable to set the tungsten content to 5.1 mass % or more. Most preferably, molybdenum is limited to 1.8 mass %, that is, it is most preferable to set the minimum of the tungsten content to 5.2 mass % or more.

[0110] Rhenium, ruthenium, and iridium elements exhibit behavior similar to tungsten. That is, these elements are gamma-forming elements that improve the creep merit index. These elements can be selectively added to the alloy. When these elements are added, the creep response of the alloy is significantly improved compared to tungsten (because the diffusivity is much slower), but the cost of the elements is high, resulting in a significant increase in cost. The addition of rhenium and iridium is preferably limited to 3.0 mass% or less, more preferably limited to 2.0 mass% or less, and most preferably limited to 1.5 mass% or less due to the cost of the elements. Ruthenium is limited to 2.0 mass%, preferably 1.5 mass% or less, because it detrimentally increases the hot cracking index (see AM1055 and AM1056).

[0111] There is a trade-off between creep resistance (from the perspective of the creep merit index), alloy stability (from the perspective of the Md number), and the corrosion resistance of the alloy (from the perspective of the chromium content). The limits of alloy stability at various levels of chromium content, determined by the equation for f(stability) (described below with reference to FIGS. 6 to 10), are shown in FIG. 5. It can be seen that as chromium increases for a given microstructure stability, the creep merit index decreases. Therefore, lowering the level of chromium is considered beneficial for the stability of the microstructure but not desirable for a high creep merit index.

[0112] The improvement in oxidation resistance, particularly corrosion resistance, is due to the addition of chromium. However, when molybdenum and tungsten are added for creep resistance and chromium is added for oxidation and corrosion resistance, the alloy has a higher tendency to form unwanted TCP phases. FIGS. 6 to 10 show the influence of the addition of tungsten and molybdenum on phase stability in alloys containing various levels of chromium. The higher the stability value, the more likely the alloy is to form TCP phases. Since the TCP phase causes deterioration of material properties over time, it is beneficial to limit or stop the precipitation of TCP phase formation. It is necessary to manage the complex trade-off between mechanical performance, oxidation / corrosion resistance, and microstructure stability.

[0113] To ensure the stability of the microstructure and avoid TCP formation, it is desirable to set the target stability number (Md), which is determined at the equilibrium temperature of 900 °C, to 0.93 or less (refer to the prior art alloys shown in Table 3). To ensure better microstructure stability and avoid TCP formation, it is more preferable to set the target stability number to 0.92 or less. From Figures 6 to 10, it can be seen that in alloys with a γ' volume fraction of 42 to 63%, the addition amounts of molybdenum, tungsten, and chromium elements follow the following formula. f(stability)=W W +W Cr +1.7W Mo

[0114] Here, f(stability) is a numerical value. When the value of f(stability) is 18.7 or less, an alloy with a stability numerical value of 0.93 or less can be obtained. Based on the formula for f(stability), when the minimum amount of tungsten is 2.0 mass%, chromium can be present in an amount of 16.7 mass% or less to satisfy the desired microstructure stability (Md is 0.93 or less). Based on the formula for f(stability), when the chromium level is 6.0 mass%, tungsten can be included in the alloy up to an upper limit of 12.7 mass%. The preferred chromium content is 8.0 mass%, and even more preferably 10 mass%. As a result, tungsten is limited to 10.7 mass% or less and 8.7 mass% or less, respectively.

[0115] When tungsten is equal to 4.0% by mass based on f(CMI), the best balance of creep resistance and corrosion resistance is achieved while maintaining alloy stability. Therefore, it is preferable to limit the maximum chromium content of the alloy to 14.7% by mass or less. Thereby, tungsten can be contained at 4.0% by mass, and the stability number Md can be made 0.93 or less. It is preferable to limit the stability number to 0.92 or less. To do this, the numerical value of f(stability) needs to be 17.8 or less. Therefore, it is preferable to limit the chromium content to 13.8% by mass or less. Thereby, the stability number is limited to 0.92 or less, and the stability of the microstructure is improved.

[0116] In combination with a high level of mechanical strength (from the perspective of creep resistance), it is necessary to limit the density of the alloy. 8.9 g / cm 3 A target density of is imposed, which is typical of the upper limit of the density of commercially used nickel-based superalloys. According to the elements within the alloy design region listed in Table 2, the tungsten element and the tantalum element have a significantly higher density than nickel and have the strongest influence on the increase in density. Figure 12 shows the influence of the tantalum element and the tungsten element on the alloy density. From Figure 12, it can be seen that the addition amounts of tungsten and tantalum need to follow the following formula. f(density) = W W + W Ta

[0117] Here, f(density) needs to be 13.9 or less in order to realize an alloy with a density of 8.9 g / cm or less. Considering the minimum required concentration of tungsten (2.0% by mass), it is necessary to limit the tantalum concentration to 11.9% by mass or less. Preferably, the density is 8.8 g / cm 3 3 ​It is limited as follows. To achieve this, the value of f(density) needs to be 11.8 or less, and thus the tantalum needs to be 9.8 mass% or less. Even more preferably, the tungsten needs to be 4.7 mass% or more, and thus the tantalum needs to be limited to up to 9.2 mass% and up to 7.1 mass%. This achieves f(density) values of 13.9 and 11.8 respectively. Even at lower levels of tantalum (e.g., 7.1 mass% or less, which is the most desirable maximum level of tantalum and includes all of the alloy examples below), the level of tungsten can be selectively increased while suppressing the density of the alloy.

[0118] The addition of carbon, boron, and zirconium provides strength to the grain boundaries. This is particularly beneficial for the creep and fatigue properties of the alloy. The carbon concentration should be in the range of 0.02 mass% to 0.35 mass%. A lower level of carbon is preferred to reduce cracking during the additive manufacturing process. Thus, it is preferably contained at 0.2 mass% or less, or 0.15 mass% or less, more preferably 0.1 mass% or less. The boron concentration should be in the range of 0.001 to 0.2 mass%. Since boron separates into the liquid phase during solidification and may cause liquation cracking during the AM process, preferably the boron concentration should be 0.03 mass% or less. More preferably, the boron concentration should be in the range of 0.02 mass% or less. The zirconium concentration must be in the range of up to 0.5 mass%, preferably 0.001 mass% to 0.5 mass%, preferably 0.05 mass% or less, or 0.035 mass% or less, more preferably 0.01 mass% or less, even more preferably 0.006 mass% or less. Preferably, the zirconium concentration is 0.005 mass% or more, more preferably 0.010 mass% or more.

[0119] When an alloy is manufactured, it is beneficial for it to be substantially free of inevitable impurities. These impurities can include the sulfur element (S), the manganese element (Mn), and the copper element (Cu). The sulfur element is preferably maintained at 0.003 mass% (30 PPM in mass conversion) or less. When sulfur is present in an amount greater than 0.003 mass%, the alloy becomes brittle and sulfur segregates at the alloy / oxide interface formed during oxidation. Therefore, the sulfur level is preferably 0.001 mass% or less. Manganese is an inevitable impurity and is limited to 0.25 mass%, preferably limited to 0.1 mass% or less. Copper (Cu) is an inevitable impurity and is preferably limited to 0.5 mass% or less. Vanadium (V) is an inevitable impurity and, since it has an adverse effect on the oxidation behavior of the alloy, is preferably limited to 0.5 mass%, preferably limited to 0.3 mass% or less, and most preferably limited to 0.1 mass% or less. This segregation can increase the peeling of the protective oxide scale. When the concentrations of these inevitable impurities exceed a predetermined level, problems arise regarding product yield, and deterioration of the material properties of the alloy is expected.

[0120] Iron exhibits behavior similar to nickel and can be added as a low-cost alternative to nickel. Furthermore, by allowing the addition of iron, the ability of the alloy to be manufactured from recycled materials is improved. Therefore, it is preferable for iron to be present in an amount of at least 0.1 mass%. However, in order to significantly reduce costs, iron can be added up to 4.0 mass%. Preferably, in order to reduce the tendency to form the undesirable Laves phase that degrades the mechanical properties of the alloy, the addition of iron is 2.0 mass% or less. Most preferably, the addition of iron is limited to 1 mass%. This results in the production of an alloy having a good ability to be recycled without compromising material performance.

[0121] To constrain the inevitable impurities in the alloy and to impart strength, it is beneficial to add hafnium (Hf) up to 1.0 mass%, preferably up to 0.5 mass%, more preferably up to 0.4 mass%, and even more preferably up to 0.2 mass%. Since hafnium is a strong carbide former, it can result in further grain boundary strengthening. Furthermore, hafnium is beneficial for improving the adhesion of the protective Al 2 O 3 oxide layer. Therefore, it is desirable to set the minimum amount of hafnium at 0.1 mass% or more, and from the viewpoint of increasing strength at the expense of increased cost, it is even more desirable to set the minimum amount of hafnium at 0.15 mass% or more.

[0122] So-called "reactive elements" (yttrium (Y), lanthanum (La), and cerium (Ce)) are added at levels up to 0.1 mass%. This is beneficial for improving the adhesion of protective oxide layers such as Al 2 O 3 etc. These reactive elements can "scavenge" harmful elements such as sulfur. Sulfur segregates at the alloy-oxide interface, weakening the bond between the oxide and the substrate and resulting in oxide peeling. Magnesium (Mg) similarly exhibits behavior of "scavenging" harmful elements and can bring a beneficial effect on mechanical properties, so it can be added up to a maximum of 0.1 mass%. Silicon (Si) can be beneficially added up to 0.5 mass%. Adding silicon at a level up to 0.5 mass% to nickel-based superalloys has been shown to be beneficial for oxidation characteristics. In particular, silicon segregates at the alloy / oxide interface, improving the bonding force of the oxide to the substrate. Thereby, oxide peeling is suppressed, and as a result, oxidation resistance is improved.

[0123] Based on the description of the present invention in this section, a broad range of the present invention is listed in Table 4. Also shown in Table 4 are the preferred ranges and the most preferred ranges. Table 4 is the composition range in mass% of the new designed alloy.

[0124]

Table 4

[0125] (Examples of the present invention) [Table 5] [Table 6]

[0126] The alloy examples shown in Table 5 and Table 6 are designed by changing the numerical value of f(HCI). Table 5 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ' volume fraction, compared with the alloys listed in Table 1. Table 6 shows the phase ratio and merit index calculated by the "alloy design" software. This is the result of comparing the conventionally used nickel-based superalloy with a high γ' volume fraction (Table 1) with the newly designed nickel-based superalloy with a high γ' volume fraction compared with the alloys listed in Table 1. The content of γ' is controlled by changing the numerical value of f(SAC). The creep merit index is controlled by the presence of Mo and W. From the results, all eight alloys have strength and creep prediction equivalent to or better than those of the prior art alloys, and show a hot cracking index superior to that of the prior art alloys. Alloys AM1005 to AM1007 are designed to have a strain aging cracking index equivalent to that of CM247LC. AM1005 is designed to have the lowest risk of hot cracking, but due to its relatively low Al content, its oxidation resistance may be the lowest. In contrast, AM1007 has the highest oxidation resistance but is designed to have the minimum hot cracking resistance. AM1006 represents the intermediate point between the two. Alloys AM1010 and AM1011 are homologous to AM1005 to AM1007 but are at the lower limit of the allowable range of f(SAC), while AM1013 is at the upper limit of the allowable range of f(SAC).

[0127] [Table 7]

Table 8

[0128] Table 7 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 8 shows the phase fractions and merit indices in the alloys listed in Table 7, calculated by the “alloy design” software. Tables 7 and 8 show the characteristics of alloys AM1018 to AM1022, which are variations of alloy AM1006. In these alloys, the ratio of Ta to Nb has been changed so as to keep the values of f(HCI) and f(SAC) constant. Ta-rich alloys have been shown to have higher strength and creep merit indices and more favorable hot cracking indices. However, Nb-rich alloys have a narrower shell freezing range and lower density and cost.

[0129]

Table 9

Table 10

[0130] Table 9 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 10 shows the phase fractions and merit indices in the alloys listed in Table 9, calculated by the “alloy design” software. Tables 9 and 10 show the composition and characteristics of alloys AM1023 to AM1028. These alloys are variations of AM1005 and use cobalt instead of nickel. Low-cobalt alloys are desirable for applications where cost is restricted and can have a narrower solidification range. On the other hand, low-cobalt alloys show lower strength merit indices and higher hot cracking indices and are thus less desirable for applications restricted by strength and / or printability. Increasing the Co content also acts to slightly suppress the formation of γ´, which is also beneficial for print quality. It is also possible to lower the cobalt level at the expense of strength.

[0131]

Table 11

Table 12

[0132] Table 11 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 12 shows the phase fractions and merit indices in the alloys listed in Table 11, calculated with "Alloy Design" software. Tables 11 and 12 show alloys AM1033 - AM1041 that contain iron (1 - 3 mass %, used in various combinations instead of cobalt and nickel). Iron is generally beneficial with respect to mechanical properties, and the main reason for its addition is to reduce costs. This is particularly seen when used instead of cobalt.

[0133]

Table 13

Table 14

[0134] Table 13 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 14 shows the phase fractions and merit indices in the alloys listed in Table 13, calculated with "Alloy Design" software. Tables 13 and 14 show a series of alloys that use Mo and W instead of Cr to maintain a certain stability merit index. Apart from high corrosion resistance, high Cr alloys have a low density and are advantageous in terms of the hot cracking index and solidification range. This can make the AM process more acceptable. They are desirable for applications where high corrosion resistance is required. In contrast, low Cr alloys have high strength and creep merit indices and are more useful for high-load applications.

[0135]

Table 15

Table 16

[0136] Table 15 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 16 shows the phase fractions and merit indices in the alloys listed in Table 15, calculated by the “Alloy Design” software. Tables 15 and 16 show alloys with Ru and Re added instead of W. All of these alloys exhibit particularly high creep merit indices, but are significantly more costly than the aforementioned alloys. Therefore, these alloys are suitable for applications that require strong creep resistance among those without cost constraints. Ru-containing alloys affect the freezing range due to Ru and have a high hot cracking index, so care must be taken before applying Ru-containing alloys.

[0137]

Table 17

Table 18

[0138] Table 17 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 18 shows the phase fractions and merit indices in the alloys listed in Table 17, calculated by the “Alloy Design” software. Tables 17 and 18 show a series of derivative alloys of AM1006 in which Nb and Ta are used instead of Ti up to the tolerance limit. In AM1060~AM1062 with a reduced Ta content, a clear decrease in density and cost is observed. However, the strength and printability decrease accordingly. A similar trend is observed in alloys AM1063~AM1065 with a reduced Nb content, but the difference is slight.

[0139]

Table 19

Table 20

[0140] Table 19 shows the nominal composition in mass % of a newly designed nickel-based superalloy with a high γ´ volume fraction. Table 20 shows the phase fractions and merit indices in the alloys listed in Table 19, calculated with "Alloy Design" software. Tables 19 and 20 show a series of derivative alloys of AM1006 with various contents of Hf and Zr, designed to improve grain boundary strength. The presence of Hf provides these alloys with excellent oxidation resistance (especially AM1067 and AM1069, where the Hf content is high). This is because Hf has a beneficial effect on the stability of the alumina scale. However, both of these elements act to draw out the terminal freezing range, increasing the hot cracking index that justifies conservative levels. In other examples, hafnium and zirconium are not added, but when hafnium and zirconium are added, a similar trend can be expected, along with improved grain boundary strength and, in the case of hafnium, improved oxidation resistance. Hafnium can be added in the absence of zirconium, and vice versa.

[0141] Compared with the baseline alloy AM1005, as the hot cracking index HCI is increased towards a limit of 1.5 (AM1006 and AM1007), it can be seen that the cracking susceptibility during AM processing increases. From Figure 14, it can be seen that when HCI is below 1.5, little or no cracking occurs during AM manufacturing, while when HCI is above 1.5 (the prior art alloys IN738, CM247C, and IN713), significant cracking occurs during AM manufacturing. The samples in Figure 14 are 10 mm × 10 mm × 10 mm cube samples, manufactured using the selective laser melting powder bed method, and evaluated for the ease of processing by additive manufacturing of selected nickel-based superalloys. The samples were manufactured under the same conditions from argon gas atomized alloy powder (15 - 53 μm) processed under an argon atmosphere (<0.1% O 2 ). The fixed energy density (2.2 J / mm 2 ) and a layer thickness of 30 μm were selected to best represent how the machine operates during manufacturing. Metallographic samples were prepared by cutting perpendicular to the build direction after the final 1 μm diamond polish to expose the X - Y plane. Figure 14 shows a typical micrograph of an optical microscope after applying a threshold derived from a histogram.

Claims

1. 4.0 to 5.6 mass% aluminum, 0.0 to 1.0 mass% titanium, 0.0 to 4.0 mass% niobium, 0.0 to 11.9 mass% tantalum, 2.0 to 12.7 mass% tungsten, 0.0 to 3.0 mass% molybdenum, 0.0 to 22.0 mass% cobalt, 6.0 to 16.7 mass% chromium, 0.02 to 0.35 mass% carbon, 0.001 to 0.2 mass% boron, 0.00 to 0.5 mass% zirconium, 0.0 to 3.0 mass% rhenium, The alloy contains 0.0 to 2.0 mass% ruthenium, 0.0 to 3.0 mass% iridium, 0.0 to 0.5 mass% vanadium, 0.0 to 0.003 mass% sulfur, 0.0 to 0.25 mass% manganese, 0.0 to 0.1 mass% magnesium, 0.0 to 4.0 mass% iron, 0.0 to 0.5 mass% copper, 0.0 to 1.0 mass% hafnium, and the balance being nickel and unavoidable impurities. The mass percentages of niobium, tantalum, and tungsten contained in the alloy are respectively W. Ｎｂ , W Ｔａ , W Ｗ A nickel-based alloy composition that satisfies the following formula: 1.1≦0.3W Ｎｂ +0.15W Ｔａ W Ｗ +W Ｔａ ≦11.8

2. The mass percentages of niobium and tantalum contained in the alloy are W Ｎｂ , W Ｔａ The nickel-based alloy composition according to claim 1, which satisfies the following formula: 1.15≦0.3W Ｎｂ +0.15W Ｔａ

3. The mass percentages of aluminum, titanium, niobium and tantalum contained in the alloy are W Ａｌ , W Ｔｉ , W Ｎｂ and W Ｔａ The nickel-based alloy composition according to claim 1 or 2, which satisfies the following formula: 5.6≦W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ ≦7.0

4. 4. The nickel-based alloy composition of claim 1, wherein the volume fraction of the gamma prime phase at 900° C. is less than or equal to 63%.

5. 5. The nickel-based alloy composition of claim 1, wherein the volume fraction of the gamma prime phase at 900°C is 42% or greater.

6. The mass percentages of tungsten and molybdenum contained in the alloy are W Ｗ , W Ｍｏ The nickel-based alloy composition according to any one of claims 1 to 5, which satisfies the following formula: W Ｗ +0.65W Mo ≧4.0

7. The nickel-based alloy composition according to claim 1 , comprising, by weight percent, 8.0% or more of chromium.

8. 8. The nickel-based alloy composition of claim 1 , comprising, by weight, up to 14.7% chromium.

9. 9. The nickel-based alloy composition of claim 1 , comprising, by weight percent, tantalum in an amount up to 9.8%.

10. 10. The nickel-based alloy composition of claim 1 , comprising, by weight percent, up to 2.0% molybdenum.

11. 11. The nickel-based alloy composition of claim 1 , comprising, by weight percent, titanium up to 0.5%.

12. 12. The nickel-based alloy composition of claim 1 , comprising, by weight, up to 0.5% hafnium.

13. 13. The nickel-based alloy composition of claim 1 , comprising up to 10.7% tungsten, by weight.

14. 14. The nickel-based alloy composition of claim 1, comprising up to 3.0% niobium, by weight.

15. 15. The nickel-based alloy composition of claim 1, further comprising, by weight percent, up to 0.5% of at least one of platinum and palladium.

16. 16. The nickel-based alloy composition of claim 1 , comprising at least 4.5% aluminum, by weight.

17. 17. The nickel-based alloy composition of claim 1, comprising, by weight percent, up to 5.3% aluminum.

18. 18. The nickel-based alloy composition of claim 1 , comprising up to 19.0% cobalt, by weight.

19. 19. The nickel-based alloy composition of claim 1 , comprising, by weight percent, up to 0.3% vanadium.

20. 20. The nickel-based alloy composition of claim 1, comprising, by weight percent, tantalum in an amount of at least 1.1%.

21. 21. The nickel-based alloy composition of claim 1 , comprising, by weight percent, tungsten in an amount of at least 2.7%.

22. The mass percentages of niobium, tantalum, titanium, platinum, palladium and aluminum contained in the alloy are W Ｎｂ , W Ｔａ , W Ｔｉ , W Ｐｔ , W Ｐｄ and W Ａｌ The nickel-based alloy composition according to any one of claims 1 to 21, which satisfies the following formula: W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ +0.125W Ｐｔ +0.225W Ｐｄ ≦7.0

23. The mass percentages of chromium, molybdenum, and tungsten contained in the alloy are W Ｃｒ , W Ｍｏ and W Ｗ The nickel-based alloy composition according to any one of claims 1 to 22, which satisfies the following formula: W Ｗ +W Ｃｒ +1.7W Ｍｏ ≦18.7

24. 24. The nickel-based alloy composition of claim 1, comprising, by weight percent, at least 8.0% cobalt.

25. 25. The nickel-based alloy composition of claim 1 , comprising, by weight, at least 0.05% niobium.

26. 26. The nickel-based alloy composition of claim 1 , comprising, by weight, at least 0.1% titanium.

27. 27. The nickel-based alloy composition of claim 1, comprising, by weight, at least 0.5% molybdenum.

28. 3. The nickel-based alloy composition of claim 1, comprising up to 7.1 wt.% tantalum and up to 5.3 wt.% aluminum.

29. 29. The nickel-based alloy composition of claim 1 , comprising, by weight percent, at least 0.1% hafnium.

30. 30. The nickel-based alloy composition of claim 1, comprising up to 0.05% zirconium, by weight.

31. 31. The nickel-based alloy composition of claim 1 , comprising, by weight percent, at least 0.005% zirconium.

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