Nickel-Based Alloys

Abstract

1.5 to 4.5 mass% aluminum, 1.1 to 3.4 mass% titanium, 0.0 to 4.0 mass% niobium, 0.0 to 5.2 mass% tantalum, 0.9 to 6.6 mass% tungsten, 0.0 to 3.0 mass% molybdenum, 0.0 to 24.0 mass% cobalt, 12.5 to 20.6 mass% chromium, 0.02 to 0.15 mass% carbon, 0.001 to 0.015 mass% boron, 0.0 to 0.1 mass% zirconium, 0.0 to 3.0 mass% rhenium, 0.0 to 2.0 mass% ruthenium, 0.0 to 3.0 mass% iridium, 0.0 to 0.5 mass% vanadium, 0. The alloy contains 0 to 1.0 mass% palladium, 0.0 to 1.0 mass% platinum, 0.0 to 0.5 mass% silicon, 0.0 to 0.1 mass% yttrium, 0.0 to 0.1 mass% lanthanum, 0.0 to 0.1 mass% cerium, 0.0 to 0.003 mass% sulfur, 0.0 to 0.25 mass% manganese, 0.0 to 0.1 mass% magnesium, 0.0 to 5.0 mass% iron, 0.0 to 0.5 mass% copper, 0.0 to 1.0 mass% hafnium, and the balance is nickel and unavoidable impurities, and the mass percentages of aluminum, titanium, niobium, tantalum, and tungsten contained in the alloy are respectively W. Al , W Ti , W Nb , W Ta and W W A nickel-based alloy composition that satisfies the following formula: 0.65≦0.3W Nb +0.15W Ta 3.6≦W Al +0.5W Ti +0.3W Nb +0.15W Ta ≦5.7 W Ta +0.92W W ≦6.1

Description

[Technical field]

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

[0002] Currently, there is a trend to transfer nickel-based superalloys, which have been successfully produced in cast or wrought forms, to AM processes, but this has proven largely inappropriate, since many of the material properties required to facilitate processing in AM processes are not met in the aforementioned alloys, resulting in materials that are significantly more difficult to process and that do not have the expected structural integrity.

[0003] In particular, it has been a great challenge to develop alloys for additive manufacturing processes that have an optimal trade-off between high strength, low density, and resistance to hot tearing. Typically, high strength alloys are processed by investment casting. Examples of common alloys used in investment casting processes are listed in Table 1, which lists the nominal composition, in weight percent, of conventional chromia-forming superalloys. In each alloy, the nickel element is the balance.

[0004] The alloys listed in Table 1 have been extensively studied for additive manufacturing. These alloys have been shown to limit certain failure mechanisms, although they may be difficult to weld. For example, these alloys are susceptible to strain age cracking. Careful control of the AM conditions during processing (scan planning, heat input, etc.) and post-processing conditions (heat treatment window and heating rate) can reduce the risk associated with strain age cracking. Similarly, resistance to strain age cracking can be improved by designing the part geometry to reduce the effect of stress concentration features such as notches. However, it is still desirable to reduce the propensity for strain age cracking in order to allow more design freedom in terms of processing conditions and geometry.

[0005] Strain age cracking is caused by two main factors: residual strain and gamma-prime precipitation. This is sought in the present invention from a compositional point of view by minimizing the content of gamma-prime elements within the limits imposed by the property targets. Illston (Patent No. US9352421B2) shows that process control (particularly the use of thin powder layers and deliberately overlapping laser scans) can minimize the accumulation of residual strain, thus improving the printability of high gamma-prime superalloys. Also, Etter et al (Patent No. US9670572B2) show that a particularly rapid ramp up to the stress relief heat treatment temperature after AM can reduce the risk of strain age cracking.

[0006] Another mechanism that cannot be easily reduced by process optimization is hot tearing, which occurs during the final stages of the solidification process and is highly dependent on the alloy chemistry. Summary of the Invention [Problem to be solved by the invention]

[0007] The objective of this invention is to develop a creep resistant, gamma prime strengthened alloy that is resistant to hot tearing and strain age cracking mechanisms by tailoring the alloy chemistry to overcome these undesirable damage mechanisms while maintaining the alloy density in the correct range and possessing competitive strength.

[0008] Combined with this improved workability, the alloys of the present invention preferably have very high levels of oxidation and corrosion resistance, which can be achieved by providing a sufficiently high ratio of chromium to titanium to form a stable and continuous protective chromia scale.

[0009] [Table 1] [Means for solving the problem]

[0010] According to the present invention, the composition contains 1.5 to 4.5 mass% aluminum, 1.1 to 3.4 mass% titanium, 0.0 to 4.0 mass% niobium, 0.0 to 5.2 mass% tantalum, 0.9 to 6.6 mass% tungsten, 0.0 to 3.0 mass% molybdenum, 0.0 to 24.0 mass% cobalt, 12.5 to 20.6 mass% chromium, 0.02 to 0.15 mass% carbon, 0.001 to 0.015 mass% boron, 0.0 to 0.1 mass% zirconium, 0.0 to 3.0 mass% rhenium, 0.0 to 2.0 mass% ruthenium, 0.0 to 3.0 mass% iridium, 0.0 to 0.5 mass% vanadium, The alloy contains 0.0 to 1.0 mass% of palladium, 0.0 to 1.0 mass% of platinum, 0.0 to 0.5 mass% of silicon, 0.0 to 0.1 mass% of yttrium, 0.0 to 0.1 mass% of lanthanum, 0.0 to 0.1 mass% of cerium, 0.0 to 0.003 mass% of sulfur, 0.0 to 0.25 mass% of manganese, 0.0 to 0.1 mass% of magnesium, 0.0 to 5.0 mass% of iron, 0.0 to 0.5 mass% of copper, 0.0 to 1.0 mass% of hafnium, and the balance being nickel and unavoidable impurities, and the mass percentages of aluminum, titanium, niobium, tantalum, and tungsten contained in the alloy are respectively W and 0.0 to 0.1 mass%. Al , W Ti , W Nb , W Ta and W W Then, there is provided a nickel-based alloy composition that satisfies the following formula: 0.65≦0.3W Nb +0.15W Ta 3.6≦W Al +0.5W Ti +0.3W Nb +0.15W Ta <= 5.7 W Ta +0.92W W <= 6.1 Such alloys have a good balance of strength, resistance to hot tearing and strain age cracking, and reasonable density.

[0011] One embodiment is a nickel-based alloy composition according to claim 1, wherein the weight percent of aluminum, titanium, niobium and tantalum in the alloy is W. Al, W Ti , W Nb and W Ta Then, the following formula is satisfied. 4.1≦W Al +0.5W Ti +0.3W Nb +0.15W Ta Such alloys have improved strength. In one embodiment, the nickel-based alloy composition comprises a nickel-based alloy having a weight percent of aluminum, titanium, niobium, and tantalum in the alloy of W. Al , W Ti , W Nb and W Ta Then, the following formula is satisfied. W Al +0.5W Ti +0.3W Nb +0.15W Ta <= 5.5 Preferably, the following formula is satisfied: W Al +0.5W Ti +0.3W Nb +0.15W Ta <= 5.4 More preferably, the following formula is satisfied. W Al +0.5W Ti +0.3W Nb +0.15W Ta <= 5.0 Such alloys have improved resistance to strain age cracking.

[0012] In one embodiment, the nickel-base alloy composition comprises a gamma prime volume fraction of 44% or less at 900° C., preferably a gamma prime volume fraction of 40% or less at 900° C., more preferably a gamma prime volume fraction of 35% or less at 900° C. and / or a gamma prime volume fraction of 19% or more at 900° C. Such alloys have high creep strength and resistance to strain age cracking.

[0013] In one embodiment, the nickel-based alloy composition comprises a nickel-based alloy having a weight percent of tungsten and molybdenum of 0.1 to 0.5 wt. W , WMo Then, the following formula is satisfied. W W +0.77W Mo ≧3.0 Preferably, the following formula is satisfied: W W +0.77W Mo ≧4.6 More preferably, the following formula is satisfied. W W +0.77W Mo ≧6.2 Such alloys have improved creep resistance.

[0014] In one embodiment, the nickel-based alloy composition has a chromium content, by weight percent, of at least 16.0%, which provides improved oxidation resistance to such alloys.

[0015] In one embodiment, the nickel-based alloy composition has chromium, by weight percent, not greater than 19.5%, preferably not greater than 19.0%, and more preferably not greater than 17.3%, which provides improved resistance to the formation of TCP phases.

[0016] In one embodiment, the nickel-based alloy composition comprises up to 4.2% tantalum, by weight. Such alloys have reduced density.

[0017] In one embodiment, the nickel-based alloy composition comprises, by weight, not more than 2.9%, preferably not more than 2.5%, molybdenum, which reduces the tendency of such alloys to precipitate deleterious TCP phases.

[0018] In one embodiment, the nickel-based alloy composition comprises, by weight, 3.0% or less of titanium, preferably 2.7% or less of titanium, which provides improved oxidation resistance due to the enhanced protection of the chromia layer.

[0019] In one embodiment, the nickel-based alloy composition comprises, by weight, 5.5% or less of tungsten, preferably 5.0% or less of tungsten, more preferably 4.5% or less of tungsten, and most preferably 3.3% or less of tungsten, resulting in a reduced density of such alloys.

[0020] In one embodiment, the nickel-based alloy composition comprises up to 3.0% niobium, by weight, which provides improved oxidation resistance.

[0021] In one embodiment, the nickel-based alloy composition contains 0.5% or less platinum and / or palladium, by weight percent, more preferably 0.1% or less, which provides a better balance between cost and improved corrosion resistance.

[0022] In one embodiment, the nickel-based alloy composition comprises at least 1.7% aluminum, by weight, to ensure a stable gamma prime phase.

[0023] In one embodiment, the nickel-based alloy composition comprises, by weight, 3.9% or less aluminum, preferably 3.7% or less aluminum, and more preferably 3.1% or less aluminum. Such alloys have reduced density.

[0024] In one embodiment, the nickel-base alloy composition comprises up to 20.0% cobalt, by weight percent, which reduces the freezing range of such alloys and therefore reduces the tendency for hot tearing.

[0025] In one embodiment, the nickel-based alloy composition comprises, by weight, tantalum in an amount greater than or equal to 0.6%, preferably greater than or equal to 1.0%, such alloys have a reduced tendency to hot tearing.

[0026] In one embodiment, the nickel-based alloy composition has, by weight, tungsten at least 1.2%, preferably at least 2.3%, and more preferably at least 2.7%. Such alloys have improved creep resistance.

[0027] In one embodiment, the nickel-based alloy composition comprises at least 5.0% cobalt, preferably at least 10.0% cobalt, by weight percent, such alloys exhibiting higher intermediate temperature creep resistance and thermal conductivity.

[0028] In one embodiment, the nickel-base alloy composition comprises, by weight, niobium at least 0.6%, preferably at least 1.2%, more preferably at least 1.60%, and even more preferably at least 1.65%. In such an alloy, better hot-tearing resistance is achieved without affecting density.

[0029] In one embodiment, the nickel-based alloy composition has, by weight, 3.8% or less of iron, preferably 2.0% or less of iron, which reduces the tendency of such alloys to form Laves phases.

[0030] In one embodiment, the nickel-based alloy composition comprises tungsten, molybdenum, and chromium in weight percent, respectively, W. W , W Mo and W Cr Then, the following formula is satisfied. W W +0.74W Mo +0.93W Cr <= 22.1 Preferably, the following formula is satisfied: W W +0.74W Mo +0.93W Cr <=21.5 Such alloys have improved stability against TCP phase formation.

[0031] In one embodiment, the nickel-based alloy composition comprises niobium, tantalum, titanium, and aluminum in weight percent, respectively. Nb , W Ta , W Ti and W Al Then, the following formula is satisfied. (0.5W Ti +0.3W Nb +0.15W Ta ) / W Al ≦1.5 In such alloys, the gamma prime stability is improved.

[0032] In one embodiment, the nickel-based alloy composition comprises a mixture of chromium and titanium in weight percent, and ... respectively. Cr and W Ti Then, the following formula is satisfied. W Cr / W Ti ≧6.0 Preferably, the following formula is satisfied: W Cr / W Ti ≧6.9 More preferably, the following formula is satisfied. W Cr / W Ti ≧7.5 Such alloys have improved oxidation resistance.

[0033] In one embodiment, the nickel-based alloy composition comprises tantalum and tungsten in weight percent, respectively, W. Ta and W W Then, the following equation is satisfied: W Ta +0.92W W <= 5.4 Preferably, the following formula is satisfied: W Ta +0.92W W <= 5.1 More preferably, the following formula is satisfied. W Ta +0.92W W ≦4.8 Even more preferably, the following formula is satisfied: W Ta +0.92W W <=4.0 These alloys have a reduced density.

[0034] In one embodiment, the nickel-based alloy composition comprises niobium and tantalum in weight percent of W. Nb and W Ta Then, the following formula is satisfied. 0.68≦0.3W Nb +0.15W Ta Preferably, the following formula is satisfied: 0.75≦0.3W Nb +0.15W Ta More preferably, the following formula is satisfied. 1.15≦0.3W Nb +0.15W Ta Such alloys have improved resistance to hot tearing. [Brief description of the drawings]

[0035] [Figure 1] Figure 1 shows the calculated strain age cracking index and hot cracking index for several commercially used chromia-forming superalloys (including those listed in Table 1). Figure 1 also shows creep resistance, strain age cracking and hot cracking limitations. The areas of interest for the present invention are shaded. [Diagram 2] Figure 2 is a contour plot of the strain age cracking index for a fixed titanium level of 0.0 wt.%. The contour plot includes isolines (desirable hot cracking index (1.6) and most favorable hot cracking index (1.2)) superimposed with lines showing desirable strengths (950 MPa and 1150 MPa). [Diagram 3] Figure 3 is a contour plot of the strain age cracking index for a fixed titanium level of 1.0 wt.%. The contour plot includes overlaid contour lines (desirable hot cracking index (1.6) and most preferred hot cracking index (1.2)), lines showing desirable strengths (950 MPa and 1150 MPa), and a line showing gamma prime stability for the right side of the graph. [Figure 4] Figure 4 is a contour plot of the strain age cracking index for a fixed titanium level of 2.0 wt.%. The contour plot includes overlaid contour lines (desirable hot cracking index (1.6) and most preferred hot cracking index (1.2)), lines showing desirable strengths (950 MPa and 1150 MPa), and a line showing gamma prime stability for the right side of the graph. [Diagram 5]Figure 5 is a contour plot of the strain age cracking index for a fixed titanium level of 3.0 wt.%. The contour plot includes overlaid contour lines (desirable hot cracking index (1.6) and most preferred hot cracking index (1.2)), lines showing desirable strengths (950 MPa and 1150 MPa), and a line showing gamma prime stability for the right side of the graph. [Figure 6] 6 shows a contour plot of the constant hot cracking index at Ti=0.0 wt.%. Superimposed on the contour plot are the preferred limits of the strain age cracking index and the desired strengths (950 MPa and 1150 MPa). [Figure 7] 7 shows a contour plot of constant hot cracking index at Ti=1.0 wt.%. Superimposed on the contour plot are the preferred limits of strain age cracking index, desirable strength (950 MPa and 1150 MPa) and gamma prime stability. [Figure 8] Figure 8 shows a contour plot of constant hot cracking index at Ti = 2.0 wt%. Superimposed on the contour plot are the preferred limits of strain age cracking index, desirable strength (950 MPa and 1150 MPa) and gamma prime stability. [Figure 9] 9 is a contour plot of constant hot cracking index at Ti=3.0 wt.%. Superimposed on the contour plot are the preferred limits of strain age cracking index, desirable strength (950 MPa and 1150 MPa), and gamma prime stability. [Figure 10] Figure 10 is a contour plot of predicted temperature at 1000 hour creep life as a function of gamma prime fraction and creep figure of merit. Overlaid on the contour plot are the commercial superalloys listed in Table 1, their gamma prime content limits, and their creep figure of merit limits. The target region of the present invention is shaded. [Figure 11] Figure 11 shows a contour plot of constant creep figure of merit (labeled values ​​multiplied by 10-15) as a function of Mo and W content. Superimposed on the contour plot are contours for fixed Cr levels of 12, 15, and 18 wt.%, with the stability number of the alloy at the desired value (Md=0.93). [Figure 12] 12 is a contour plot of the Cr / Ti ratio as a function of Cr content and Ti content, plotting the commercial superalloys listed in Table 1. The target region of the present invention is shaded. [Figure 13] 13 is a contour plot showing the gamma matrix stability in terms of Md number according to the present invention as a function of Cr content and W+0.77Mo (wt%), with a horizontal line superimposed on the contour plot corresponding to the desired creep resistance. [Figure 14] FIG. 14 is a contour plot showing the predicted non-equilibrium solidification temperature range of the alloy in the design domain as a function of Co content and strain age cracking index. [Figure 15] FIG. 15 is a contour plot of the calculated alloy density in the current design space as a function of Ta and W contents in wt. %. [Figure 16] 16 shows photomicrographs of the novel nickel-base superalloy of the present invention, ABD-900AM, and conventional superalloys, Haynes 282 and IN738, which show a lower microcrack density in the present invention. [Figure 17] Figure 17 is a micrograph after heat treatment at 760°C for 1000 hours, showing the effect of increasing molybdenum content on the TCP phase content. DETAILED DESCRIPTION OF THE PREFERRED EMBODIMENTS

[0036] Traditionally, nickel-base superalloys have been designed on an empirical basis. Thus, the chemical composition of a nickel-base superalloy has been identified through time-consuming and expensive experimental development involving small-scale processing of limited amounts of material and subsequent characterization of behavior. The alloy composition found to exhibit the best, or most desirable, combination of properties is then adopted. The large number of alloying element families that can achieve this combination indicates that these alloys are not fully optimized and that improved alloys are likely to exist.

[0037] In superalloys, chromium (Cr) and aluminum (Al) are typically added to provide oxidation / corrosion resistance, and cobalt (Co) is added to improve resistance to sulfidation. For creep resistance, molybdenum (Mo), tungsten (W) and cobalt are added because they inhibit thermally activated processes (e.g. dislocation climb) that determine the rate of creep deformation. For static and cyclic strength, aluminum (Al), tantalum (Ta), niobium (Nb) and titanium (Ti) are added because they promote the formation of the precipitation hardening phase gamma prime (γ´), which is coherent with the face-centered cubic (FCC) matrix phase called gamma (γ).

[0038] The model-based approach used to identify new grades of nickel-base superalloys is described herein by the term "Alloy Design" (ABD) methodology. This approach utilizes a framework of computational materials models to estimate design-relevant properties over a very wide compositional range. In principle, this alloy design tool allows the so-called inverse problem to be solved; that is, the optimum alloy composition that best satisfies specified design constraints can be identified.

[0039] The first step in the design process is to define the table of elements and the associated upper and lower compositional limitations. In the present invention, the compositional limitations for each element are considered when adding each element, called the "Alloy Design Space". These compositional limitations are detailed in Table 2. Table 2 shows the alloy design space, in weight percent, that was explored using the "Alloy Design" method.

[0040] [Table 2]

[0041] The balance is nickel. Carbon, boron and zirconium levels were fixed at 0.04%, 0.005% and 0.01%, respectively.

[0042] The second step is based on thermodynamic calculations to calculate the phase diagrams and thermodynamic properties of a particular alloy composition. This is often called the CALculation of PHAse Diagrams (CALPHAD method). These calculations are performed at the typical service temperature of the new alloy (900°C) to obtain information about the phase equilibria (microstructure).

[0043] The third step involves identifying an alloy composition with the desired microstructure. For nickel-base superalloys that require good resistance to creep deformation, creep rupture life improves gradually as the volume fraction of the precipitation hardening phase γ' increases. The range of γ' volume fraction where creep rupture life is most beneficial is 60-70%. These high levels of γ' can adversely affect alloy workability. Therefore, in some cases, γ' must be lowered to manage this tradeoff. A decrease in creep resistance is observed above 70% γ' volume fraction.

[0044] Also, the gamma / gamma' lattice mismatch must be positive or negative, whichever is smaller, to lose coherency. Therefore, the limit depends on the absolute value. The lattice mismatch δ is defined as the mismatch between the gamma and gamma' phases and is given by the following formula:

[0045]

number

[0046] Here, α γ and α γ´ are the lattice constants of the γ and γ′ phases.

[0047] Alloys based on inappropriate microstructures are also ruled out by estimates of susceptibility to morphologically close packed (TCP) phases. Using CALPHAD modeling in the present calculations, the formation of the deleterious TCP phases sigma (σ), P, and mu (μ) is predicted.

[0048] The model thus identifies all compositions within the design space that have a desired calculated volume fraction of γ', where the lattice mismatch of γ' is less than a given absolute value and the total volume fraction of TCP phases is less than a given magnitude.

[0049] In the fourth step, figures of merit are estimated for the identified alloy compositions remaining in the data set. Example figures of merit include creep figure of merit (which indicates an alloy's creep resistance based on average composition only), strength figure of merit (which indicates an alloy's precipitation yield strength based on average composition only), solid solution figure of merit (which indicates an alloy's solid solution yield strength based on average composition only), density, and cost.

[0050] In the fifth stage, the calculated figures of merit are compared to constraints on the desired behavior, and these design constraints are considered as boundary conditions for the problem. All compositions that do not satisfy the boundary conditions are rejected. At this stage, the size of the test data set is made very small.

[0051] The sixth and final step involves analyzing the remaining composition dataset. This can be done in a variety of ways. One way is to sort through the database for alloys with the highest figures of merit, such as the lightest alloys, the most creep resistant alloys, the most oxidation resistant alloys, and the least expensive alloys. Alternatively, the database can be used to determine the relative tradeoffs in performance that result from different combinations of properties.

[0052] Explain seven examples of merit figures.

[0053] The first figure of merit is the creep figure of merit. The most important observation is that the time-dependent deformation (i.e. creep) of nickel-base superalloys occurs by dislocation creep with initial activity limited to the gamma phase. Thus, dislocation segments are rapidly pinned at the gamma / gamma interface due to the large fraction of gamma' phase. The rate-limiting step is the escape of the trapped configuration of dislocations from the gamma / gamma' interface. It depends on the local chemistry (in this case the composition of the gamma phase) which causes the critical effect of the alloy composition on the creep properties.

[0054] The physics-based microstructural model assumes that the loading is uniaxial. <001> In the case where the creep strain ε is along the crystallographic direction, · The set equation is:

[0055]

number

[0056] 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 T 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.

[0057]

number

[0058] term K CF describes the proximity of cubic grains in these alloys. Equation 3 describes the dislocation multiplication process, which requires an estimate of the multiplication parameter C and the initial dislocation density. The term D eff is the effective diffusivity that controls the ascending process at the particle / matrix interface.

[0059] In the above, the composition dependence is expressed by two terms φ p and D eff Therefore, assuming the microstructure is constant (the microstructure is largely controlled by heat treatment), φ p Since is fixed, the dependence on chemical composition is D eff For the purposes of the alloy design modeling described herein, it can be appreciated that it is not necessary to perform a full integration of Equations 2 and 3 for each prototype alloy composition. Instead, the first order figure of merit M creep is used. creep is calculated using the following formula:

[0060]

number

[0061] where x i is the atomic fraction of solute i in the γ phase. D i ~ is the appropriate interdiffusion coefficient.

[0062] The second figure of merit is the strength figure of merit. For high nickel-base superalloys, most of the strength comes from the precipitate phases. Therefore, optimizing the alloy composition to maximize the precipitation strength is an important design consideration. Based on hardening theory, the strength figure of merit M strength is proposed. This index takes into account the maximum possible precipitation strength (determined as the point at which the transition of dislocation shear from weak to strong bonds occurs) and is approximated using the following equation:

[0063]

number

[0064] 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.

[0065] From equation (5), it is clear that the defect energy in the gamma-phase (e.g., the antiphase boundary APB energy) has a significant effect on the deformation behavior of nickel-base superalloys. Increasing the APB energy was found to improve the mechanical properties, including tensile strength and resistance to creep deformation. A study of the APB energy was carried out for a number of Ni-Al-X systems using density functional theory. The study calculated the effect of ternary elements on the APB energy of the gamma-phase and assumed a linear superposition of the effects of each ternary element addition when considering complex multicomponent systems. This resulted in the following equation:

[0066]

number

[0067] where x Cr , x Mo , x W , x Ta , x Nb and x Ti are the atomic percentages of chromium, molybdenum, tungsten, tantalum, niobium, and titanium in the gamma prime phase. The composition in the gamma prime phase is determined by phase equilibrium calculations.

[0068] The third figure of merit is density. Density ρ was calculated using a simple rule of mixtures and correction factors, where ρ i is the density of a given element, and x i is the atomic fraction of the alloying element.

[0069]

number

[0070] The fourth figure of merit is cost. To estimate the cost of each alloy, we applied the simple rule of mixtures, where the cost of each alloy is the mass fraction of the alloying element x i 2. Current (2016) raw material costs of alloying elementsi The value multiplied by was used.

[0071]

number

[0072] This estimate assumes that processing costs are the same for all alloys, i.e. product yields are not affected by composition.

[0073] The fifth figure of merit is based on the elimination of alloy candidates based on unsuitable microstructures made based on susceptibility to TCP phases. To do this, the d-orbital energy levels (designated Md) of the alloying elements are used to determine the total effective Md level according to the following formula:

[0074]

number

[0075] where x i represents the mole fraction of element i in the alloy. A higher value of Md indicates a higher possibility of TCP formation.

[0076] The sixth figure of merit is the strain age cracking susceptibility index. The percentage of gamma-´ in an alloy is related to its chemical composition and, most directly, to the amount of gamma-´ forming elements (aluminum, titanium, niobium, and tantalum). In turn, the higher the gamma-´ percentage, the higher the risk of strain age cracking during post-thermal treatment. This index is derived by correcting the coefficients of the heavier elements and summing the mass percentages of each element. The coefficients of the heavier elements are corrected to account for their larger atomic weights relative to that of aluminum. The strain age cracking susceptibility index is given by the following formula:

[0077]

number

[0078] Here, W Al , W Ti , W Nb and W Ta are the mass percentages of aluminum, titanium, niobium and tantalum, respectively, contained in the alloy. Lower values ​​of this strain aging index indicate a smaller potential fraction of gamma ' and therefore a lower risk of strain aging cracking, improving processability by additive manufacturing. Higher values ​​of the strain aging index indicate a larger potential fraction of gamma ' and therefore better mechanical properties in terms of strength and creep resistance.

[0079] The seventh figure of merit is based on the solidification behavior of candidate alloys as predicted by the Scheil-Gulliver model, which ranks the susceptibility to hot tearing based on composition. In this approach, the temperature range at the final stage of solidification (90-99% solids) is considered to represent the region where the alloy becomes vulnerable to hot tearing, since at this stage the liquid supply is more likely to be limited by the bridged network of solid material. The temperature range at 40-90% solids is considered to be a safe region, since at this stage the restrictions on the liquid supply are greatly relaxed. The temperature range below 40% solids is considered to be unsuitable, since the liquid is dominant.

[0080] According to the ranking system used by Clyne and Davis for casting alloys, the hot tearing index is defined as the ratio of the vulnerable temperature range to the safe temperature range as follows:

[0081]

number

[0082] The lower the value of this index, the lower the risk of hot tearing.

[0083] Using the ABD methodology described above, the alloy composition of the present invention was identified. The design intent is to develop an alloy that has excellent creep and oxidation resistance (achieved by a high gamma prime volume fraction and a stable protective chromia scale, respectively) and excellent fabricability by additive manufacturing (achieved by having excellent resistance to hot tearing mechanisms) compared to other conventional nickel-base superalloys. Additionally, properties including microstructural stability and alloy density are optimized.

[0084] The material properties (determined using the ABD method) of the conventional compositions (listed in Table 1) are listed in Table 3. The design of the new alloys was considered in relation to the predicted properties listed for these alloys. Table 3 shows the calculated phase fractions and figures of merit generated by the "Alloy Design" software. The results are for the nickel-base superalloys listed in Table 1.

[0085] The design principles of the new alloy are explained below.

[0086] [Table 3]

[0087] Alloys with high strain age cracking indexes are known to be difficult to fabricate based on the welding literature, and the same tendency is commonly observed in AM. Lowering the strain ageing figure of merit reduces the tendency of this failure mechanism, but high temperature strength (in terms of creep strength and tensile strength) and oxidation resistance (in terms of the protective oxide scale formed) depend on a relatively high content of γ´ forming elements that suppress the effects that hinder the improvement of the strain ageing index. To overcome this, strain age cracking can be managed to some extent by other methods. For example, careful control of the AM conditions during fabrication (scanning strategy, heat input, etc.) and post-processing conditions (heat treatment window and heating rate) can limit the risk of strain age cracking. Similarly, the risk of strain age cracking can be reduced by designing the part geometry to reduce stress concentration features that lead to localized strain age cracking. See, for example, US9352421B2 and US9670572B2. However, there are cases where it is undesirable to have to use such methods.

[0088] Apart from strain age cracking, another manufacturing defect that can occur during the additive manufacturing process is hot cracking. Hot cracking can occur during the solidification process. This manufacturing defect mechanism is caused by changes in the liquid composition during the solidification of the alloy. The risk of hot cracking cannot be managed in the same way as strain age cracking. Post-processing or design modifications do not affect this mechanism, although careful control of the AM conditions during processing may help to some extent. The most direct way to improve resistance to hot cracking is to change the alloy chemistry. Improving resistance to hot cracking significantly improves the processability of the alloy for additive manufacturing, even in cases where a high strain age cracking index must be tolerated. The objective of the present invention is to identify alloys (see FIG. 1) that have low values ​​of hot cracking index compared to other alloys, and a relatively high strain ageing index in the range where strain age cracking can be managed by processing conditions.

[0089] Figures 2-5 show the relationship between the amount of aluminum, niobium, and tantalum added when titanium is set at different levels. These elements are added primarily to form the γ' phase and to control the γ' volume fraction. Elements that form the γ´ phase can reduce the processability of an alloy by AM due to an increased strain age figure of merit (Eq. 10). Therefore, the combination of these elements needs to be optimized to provide the desired balance between limiting the potential for strain age cracking during AM processing and the γ´ volume fraction (providing strength in terms of creep resistance and tensile strength).

[0090] The strain age cracking exponent f(SAC) is modeled in the element ranges of the present invention based on the elements that have the strongest influence on strain age cracking, i.e., the contents of aluminum, titanium, niobium and tantalum satisfy the following constraints: f(SAC)=W Al +0.5W Ti +0.3W Nb +0.15W Ta

[0091] where f(SAC) is a number and W Al , W Ti , W Nb and W Ta In terms of strain age cracking resistance, f(SAC) is selected to be equal to or greater than that of IN738, with a value of 5.7 or less. In order to improve the processing and post-treatment of the alloy, it is preferable to lower the strain age cracking index. That is, W Al +0.5W Ti +0.3W Nb +0.15W Ta ≦5.5 or 5.4 is preferable, and W Al +0.5W Ti +0.3W Nb +0.15W Ta It is more preferred that the limit be ≦5.0. Dashed lines illustrating various limits for strain age cracking are shown in Figures 6-9.

[0092] The combination of aluminum, titanium, niobium, and tantalum must also be optimized to obtain the desired creep resistance and tensile strength. Both creep resistance and tensile strength increase with increasing gamma-proportion. To achieve the desired gamma-proportion volume fraction (19%) at 900°C, W is Al +0.5W Ti +0.3W Nb +0.15W Ta ≧3.6. This limit is defined based on the required minimum level of creep resistance (see below in connection with FIG. 10). Furthermore, it is preferable to have a γ′ fraction of 26% or more. This allows W Al +0.5W Ti +0.3W Nb +0.15W Ta A preferred value of ≧4.1 is obtained.

[0093] For f(SAC) in the range of 3.6-5.7, alloys with γ´ fractions of approximately 19-44% at 900°C are produced, providing alloys with a good combination of high creep strength and excellent resistance to strain age cracking during additive manufacturing and post-processing.

[0094] The elements platinum and palladium behave similarly to tantalum, titanium and niobium; that is, they are gamma-formers that increase the antiphase boundary energy. These elements can be selectively added to the alloy in place of tantalum, titanium and niobium. Benefits of this can include improved hot corrosion resistance. The "aluminum equivalents" of platinum and palladium require correction factors (determined from their densities relative to aluminum) of 0.125 and 0.225, respectively. However, the high cost of adding these elements may limit the amount of these elements added. Thus, these elements may be present at a maximum of 1.0 wt. % each, and are preferably limited to no more than 0.5 wt. %, and most preferably limited to no more than 0.1 wt. %. This range provides the optimum balance between cost and improved corrosion resistance. To achieve the optimum balance between creep performance and AM workability, it is preferred that the following formula be satisfied: W Al +0.5W Ti +0.3W Nb +0.15W Ta +0.125W Pt +0.225W Pd ≧3.6 and W Al +0.5W Ti +0.3W Nb +0.15W Ta +0.125W Pt +0.225W Pd <= 5.7

[0095] Here, W Pt and W Pd indicate the mass percentages of platinum and palladium contained in the alloy, respectively.

[0096] The relative proportions of aluminum, titanium, niobium and tantalum must also be balanced to minimize the risk of hot cracking during solidification in the AM process. Hot cracking is the tearing of solidified metal, usually caused by insufficient liquid supply at the later stages of solidification. That is, the solidified metal is allowed to tear open due to thermal strain. This can be mitigated by minimizing the hot cracking index, as shown in Figures 6-9. In the present invention, it is desirable to have a hot cracking index of 1.6 or less. This represents a lower risk of hot cracking than the conventional alloys Haynes282, IN792, IN738, Rene88DT and U720Li. Such hot cracking index is strongly related to the levels of niobium and tantalum, and the hot cracking index can only be estimated using the following formula in the composition range of the present invention based on the content of these alloys. f(HCI)=0.3W Nb +0.15W Ta

[0097] Here, f(HCI) is a value that must be 0.65 or more to provide a hot cracking index of 1.6 or less at Co and Cr levels close to the composition of ABD-900AM. The hot cracking index is further reduced by increasing the Co and Cr levels. By ensuring an f(HCI) of 0.68 or more, the hot cracking resistance can be further improved. By ensuring an f(HCI) of 0.75 or more, the hot cracking resistance can be further improved. More preferably, the hot cracking index is 1.2 or less, which indicates a reduced risk of hot cracking. That is, f(HCI) is more preferably 1.15 or more. As shown in FIG. 16, the alloy ABD-900AM of the present invention has a lower hot cracking index and therefore less hot cracking than the conventional alloys Haynes 282 and IN738.

[0098] Additionally, it is desirable for the alloy to contain sufficient amounts of Ti, Nb, and Ta to ensure a reasonable strength figure of merit. The strength figure of merit contours of 950 MPa and 1150 MPa are shown in Figures 6-9. With reference to Figures 6-9, it can be seen that the strength figure of merit increases with increasing titanium content. Increasing f(SAC) is also associated with increasing strength figure of merit. To achieve a strength figure of merit higher than Haynes 282 and ATI 718+ across the alloy range, a Ti content of at least 1.1 wt% is used. From Figure 7, it can be seen that the combination of the minimum value of f(SAC) (3.6), the maximum value of f(HCI) (1.6), and the minimum value of Al (1.5 wt%), when titanium is 1.0 wt%, a strength figure of merit of 950 MPa is not achieved at all levels of Al. From Figure 8, it can be seen that under the same conditions (except Ti = 2.0 wt%), a strength figure of merit of 950 MPa is easily achieved, and many of the alloy regions have strength figures of merit of 1150 MPa or higher (this is achieved when f(SAC) is in the preferred range (4.1 or higher).

[0099] Instead of aluminum, titanium, niobium and tantalum are added and these elements are distributed in the gamma-phase. If the combination of these elements (atomic percentages) is significantly greater than the total aluminum content (atomic percentages), the stability of the gamma-phase may decrease and unwanted phases such as delta (δ) and eta (η) may form. The element ratios (atomic percentages) are estimated by converting the mass percentages of titanium, niobium and tantalum by their relative densities to aluminum (using coefficients of 0.5, 0.3 and 0.15, respectively). Therefore, it is preferred that the ratio of the sum of titanium, niobium and tantalum to aluminum is kept below 1.5. That is, it is preferred that the following formula is satisfied: (0.5W Ti +0.3W Nb +0.15W Ta ) / W Al ≦1.5

[0100] Based on these requirements, the maximum level of aluminum in the present invention is 4.5 wt.%, which corresponds to the maximum allowable value of f(SAC) and the minimum value of f(HCI) at 1.1 wt.% titanium. This corresponds to f(HCI)=0.65 and W Ti This can be seen by substituting f(SAC)=1.1 into the formula for f(SAC)=5.7. More preferably, the maximum aluminum content is 3.9% or 3.7%, by weight. These levels allow for the inclusion of relatively high levels of titanium (e.g., 2.5%, by weight), thereby reducing density. More preferably, the maximum aluminum content is 3.1%, by weight, which allows for alloys with this high titanium level and f(SAC)=5.0.

[0101] The minimum allowable aluminum content in the present invention is 1.5 wt. %, based on the requirement that the gamma prime stability be 1.5 or greater and the minimum desired value of f(SAC). This can be seen by evaluating the intersection of the gamma prime stability contour line with the f(SAC)=3.6 contour line with reference to Figures 6-9. In the current design space, alloys with less than 1.5 wt. % aluminum have an unstable gamma prime stability or do not have enough gamma prime stability to achieve the creep life goal. More preferably, the minimum aluminum content is 1.7 wt. %. This increases the certainty that the alloy will be to the right of the gamma prime stability contour line shown in Figures 6-9, ensuring a stable gamma prime stability at a value of f(SAC)=4.1. The reduction in titanium content provides the alloy with good oxidation resistance. In order to reduce the titanium content, the aluminum content must be increased to maintain the gamma prime stability in the present invention within the desired range. Conversely, increasing the titanium content reduces the allowable aluminum content.

[0102] The present invention contains sufficient amounts of niobium or tantalum to minimize the risk of hot tearing, as described by the function f(HCI). As will be explained later (see FIG. 15), the tantalum content is limited in combination with tungsten to keep the alloy density low. 8.6 g / cm 3In order to reduce the density towards the target density of 1.0, the maximum allowable tantalum content is 5.2 wt.%, based on the minimum tungsten level required to ensure creep resistance (discussed below with reference to FIG. 11). Thus, to ensure a hot cracking index of 1.6 or less at this alloy density, niobium is preferably 0.6 wt.% or more. More preferably, to ensure a hot cracking index of 1.2 or less at this density, niobium is 1.2 wt.% or more. It is even more beneficial to have a niobium level of 1.60% or more or 1.65% or more to reduce the tendency to hot cracking. In order to reduce the hot cracking index, a minimum of 0.6 wt.% tantalum is preferred. To fall below the preferred hot cracking index (1.2), a tantalum content of 1.0 wt.% or more is preferred. In this case, a maximum niobium limit of 3.0 wt.% is particularly preferred (discussed below in relation to oxidation resistance).

[0103] Another feature of the alloy of the present invention is its good creep resistance. The relationship between creep figure of merit, gamma prime volume fraction, and predicted creep life is shown in Figure 10. The minimum objective of the present invention is an operating temperature of at least 20°C higher than Haynes 282. This requires a minimum gamma prime of 19% and a minimum creep figure of merit of 5.0 x 10 15 More preferably, the creep temperature capability of the present invention is at least 40°C higher than that of Haynes 282, which requires a γ´ fraction of 26%. Increasing the creep figure of merit improves resistance to diffusion-controlled creep damage and extends creep life. Therefore, the creep figure of merit should be increased to 5.5×10 15 More preferably, the creep merit index is 6.0×10 or more. 15 This makes it possible to achieve creep performance comparable to that of U720Li and IN738.

[0104] A higher gamma-percentage will result in a longer creep life due to increased resistance to dislocation control damage. However, an excessively high gamma-percentage will make the alloy difficult to process by AM (especially during post-fabrication heat treatment) due to the risk of strain age cracking. It is therefore desirable to limit the gamma-percentage to 44%, which is equivalent to IN738. This will provide creep performance based on the creep figure of merit equivalent to or greater than IN738. More preferably, the gamma-percentage is limited to 40%, which will provide creep capability equivalent to or greater than U720Li while minimizing the risk of strain age cracking. Most preferably, the gamma-percentage is limited to 35%. Such an alloy will have a creep life equivalent to Rene88DT and a minimum desirable creep figure of merit (5.0×10 15 ), further reducing the risk of strain age cracking.

[0105] The molybdenum and tungsten contents in the present invention are optimized to meet the desired creep figure of merit specified above. As shown in Figure 11, the relationship between the levels of these elements and the creep figure of merit can be estimated by the following formula: f(creep)=W W +0.77W Mo

[0106] Generally speaking, a creep merit factor of 5.0 x 10 15 To achieve the above desirable values, the function f(creep) must be 3.0 or more. A more desirable creep figure of merit (5.5×10 15 ), f(creep) must be 4.6 or more. The most favorable creep merit index (6.0×10 15), f(creep) should be 6.2 or greater. Superimposed on the graph are contour lines of certain alloy stability to TCP phase formation at various levels of chromium. Compositions to the left of each contour line have lower Md stability numbers at certain levels of chromium. This shows that the creep figure of merit can be higher at lower levels of chromium and that a trade-off must be managed between creep resistance, oxidation resistance, and stability.

[0107] Since TCP phases are very detrimental to the creep properties of the alloy, the maximum molybdenum content in the present invention is 3.0 wt.% in order to reduce the level of TCP phases while ensuring adequate resistance to hot corrosion. The balance of Mo in relation to Cr and W with respect to maintaining the stability of the alloy is explained later in connection with Figures 11 and 12. Figure 17 shows the microstructure after 1000 hours of heat treatment at 760°C. The left diagram of Figure 17 is an example of an alloy that falls outside the scope of the present invention only due to the excessive molybdenum (3.3 wt.%). The right diagram of Figure 17 is the ABD-900AM of the present invention. Figure 17 shows the importance of limiting molybdenum to 3.0 wt.% or less in the alloy of the present invention. More preferably, the maximum molybdenum content is 2.9 wt.% or even 2.5 wt.% in order to improve the resistance to hot corrosion. Based on this, the level of tungsten in the present invention is 5.0×10 tungsten or less in order to provide adequate resistance to creep damage. 15 % (in terms of creep figure of merit) is 0.9% by weight or more. More preferably, the tungsten content is 1.2% by weight or more, especially when the molybdenum content is 2.5% by weight or less. The higher the value of the creep figure of merit, the higher the value of tungsten required. 15 In order to ensure a creep figure of merit of 0.01%, the minimum tungsten content is preferably 2.3%, more preferably 2.7%, by weight, especially for Mo contents of 3.0% and 2.5%, by weight, respectively.

[0108] The elements rhenium, ruthenium and iridium behave similarly to tungsten; that is, they are gamma-forming elements that improve the creep figure of merit. These elements can be selectively added to the alloy. The addition of these elements significantly improves the creep response of the alloy compared to tungsten (due to its much slower diffusivity), but significantly increases the cost due to the high cost of the elements. Due to the element cost, the addition of rhenium is preferably limited to 3.0% or 2.0% by weight or less, and the addition of iridium is preferably limited to 3.0% by weight or less, even more preferably limited to 2.0% by weight or less, and most preferably limited to 1.5% by weight or less. Ruthenium is limited to 2.0% by weight and preferably to 1.5% by weight, because it adversely increases the hot tearing index.

[0109] The alloy of the present invention is required to have excellent oxidation and corrosion resistance at temperatures above 900° C. In this regime, the passive chromium oxide scale (Cr 2 O 3 ) is a suitable means of providing the alloy with resistance to further oxidation damage. To produce a continuous chromium scale, the alloy needs to have a significant chromium content. Based on the known properties of the prior art alloys, a minimum chromium content of 12.5% ​​by weight is desirable, which would provide oxidation resistance equivalent to that of IN792. It is desirable to increase the chromium content beyond this level, which would increase the supply of chromium atoms for scale formation, thus minimizing the vulnerable transition period during which oxides form. More preferably, the minimum chromium content in the present invention should be 16.0% by weight, which would provide oxidation resistance equivalent to or better than that of IN738.

[0110] In addition to forming a continuous chromia scale, the present invention also encompasses alloys with a rapidly forming dense chromia scale, i.e., a chromia scale that forms early in environmental exposure. Titanium is known to be detrimental to the stability of the chromia scale. Titanium in particular is detrimental because it increases the diffusivity of oxygen across the scale, damaging the underlying material in the process (see Chromia layer growth on a Ni-based superalloy: Sub-parabolic kinetics and the role of titanium, Cruchley et al. Corrosion Science 2013). Based on this, the ratio of chromium to titanium content (wt%) is used to measure the stability of the chromia scale. Based on the prior art alloys listed in Table 3, a Cr / Ti ratio of 6.0 or more is desirable (as shown graphically in FIG. 12) to achieve higher oxidation resistance than IN738, U720Li, and Rene88DT. More preferably, the Cr / Ti ratio is 6.9 or more, and more preferably 7.5 or more. This allows for the production of a more stable and continuous protective oxide scale.

[0111] Based on the maximum chromium content of the alloy (20.6%, by weight) (described below with reference to FIG. 13), it can be seen that the titanium content of the present invention is preferably equal to or less than 3.4%, by weight, when the Cr / Ti ratio is 6.0. Based on the more preferred Cr / Ti ratio (6.9), the maximum titanium content is preferably 3.0%, by weight. More preferably, the Cr / Ti ratio is 7.5, which corresponds to the preferred maximum titanium content (2.7%, by weight), which provides better resistance to oxidation damage.

[0112] The alloying addition of niobium is also known to have a detrimental effect on oxidation. This is because niobium forms grain boundary carbides that are particularly detrimental 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, the addition of niobium is beneficial for improving hot cracking resistance (discussed later with reference to Figures 6-9) and strength. Therefore, niobium is limited to 4.0 wt.%, and more preferably to 3.0 wt.% or less.

[0113] The elements molybdenum, tungsten and chromium all contribute to the tendency of an alloy to form deleterious TCP phases. The relationship between these elements and the stability index Md, which represents the susceptibility of a composition to TCP formation, is shown in Figure 13. In the present invention, it is desirable to have a stability index of 0.93 or less, based on the prior art alloy values ​​in Table 3. This constraint is modeled in the preferred alloy range as follows: f(Md)=W W +0.77W Mo +0.93W Cr

[0114] Here, if the value of f(Md) is 22.1 or less, the stability merit index is 0.93. If the value of f(creep) is 5.0×10 15 Based on the above requirement that the creep factor of merit is equal to or greater than 5.5×10, it can be seen that the maximum allowable chromium content is 20.6% by weight. Reducing the maximum chromium content to 19.5% improves stability. More preferably, the creep factor of merit is 5.5×10 15 and the function f(creep) must be 4.6 or more. This allows the Cr content to be set to a more preferable maximum value (19.0 mass%). Most preferably, the creep figure of merit is 6.0×10 15 Thus, the maximum value of Cr is most preferably 17.3 wt.%. To improve alloy stability, it is preferred to have an f(Md) level of 21.5 or less.

[0115] The risk of hot tearing in the present invention can also be mitigated by limiting the freezing range of the alloy, since this reduces the vulnerable time window during which hot tearing can occur. In the present invention, the freezing range (predicted by the Shale Gulliver approach) is preferably below 280°C, which requires a cobalt content of 24% by weight or less (see Figure 14). More preferably, the shale freezing range is below 270°C, which requires a cobalt content of 20% by weight or less.

[0116] However, cobalt is beneficial to certain material properties of nickel-base superalloys, such as intermediate temperature creep resistance (by increasing the stacking fault energy of the matrix phase) and thermal conductivity. Therefore, a minimum of 5.0% cobalt, by weight, is preferred. More preferably, the minimum cobalt content is 10.0% by weight, which provides a higher stacking fault energy than IN738.

[0117] In addition to maximizing the mechanical properties of the alloy, it is desirable to minimize density. This is accomplished by controlling the addition of the higher atomic weight elements, particularly tungsten and tantalum. The predicted alloy density versus tungsten and tantalum elemental levels, with the other elements held at average levels, is shown in Figure 15. This results in a model that is within the elemental ranges of the present invention and satisfies f(HCI) and f(SAC), as follows: f(density)=W Ta +0.92W W

[0118] Here, f(density) is a numerical value. When the value of f(density) is 6.1 or less, the maximum target value of alloy density (8.6 g / cm3) is usually reached. 3For example, lowering the level of titanium will adversely affect density, but for such alloys, higher additions of Nb and Ta will be required to meet the f(SAC) requirements, which may cause f(density) to underpredict density. Increasing the levels of other elements (e.g., molybdenum) may result in an actual density greater than that predicted by f(density). An f(density) of 5.4 or less will result in a desirable, low density alloy. More preferably, the alloy density should be less than 8.55 g / cm. 3 ≦f(density)≦5.1. This results in lighter parts (used in rotating applications) that require less creep strength. In this case, the value of f(density) is ≦5.1. Even lower density alloys can be achieved with f(density) ≦4.8 or even ≦4.0.

[0119] These limitations set upper limits on the tantalum and tungsten content: tantalum is no more than 5.2 wt.%, most preferably no more than 4.2 wt.%, (taking into account the creep-based tungsten minimum requirement (0.9 wt.%) discussed above). The maximum tungsten is no more than 6.6 wt.%, preferably no more than 5.5 wt.%, based on f(density). Based on the most preferred minimum limit for tantalum (1 wt.%), to achieve the most preferred density, the maximum tungsten content is most preferably no more than 4.5 wt.%.

[0120] Additions of carbon, boron and zirconium are necessary to provide strength to the grain boundaries. This is particularly beneficial for the creep and fatigue properties of the alloy. However, these elements can all crack during the AM process by solidification and / or liquation cracking mechanisms. Therefore, the carbon concentration should be in the range of 0.02% to 0.15% by weight. Lower levels of carbon are preferred to reduce cracking during the additive manufacturing process. Therefore, carbon is preferably contained at 0.08% by weight or less. Since boron strongly separates into the liquid phase during solidification, the boron concentration should be in the range of 0.001 to 0.015% by weight, preferably 0.008% by weight or less. Similarly, the zirconium concentration is 0.1% by weight or less, more preferably 0.0% to 0.02% by weight.

[0121] When the alloy is produced, it is beneficial that it is substantially free of unavoidable impurities. These impurities may include elemental sulfur (S), elemental manganese (Mn) and elemental copper (Cu). Elemental sulfur is preferably maintained at 0.003% by weight or less (30 PPM by weight). More than 0.003% by weight of sulfur embrittles the alloy and causes sulfur segregation at the alloy / oxide interface formed during oxidation. Therefore, the sulfur level is preferably 0.001% by weight or less. Manganese is an unavoidable impurity and is limited to 0.25% by weight, preferably 0.1% by weight or less. Copper (Cu) is an unavoidable impurity and is preferably limited to 0.5% by weight. Vanadium (V) is an unavoidable impurity and is preferably limited to 0.5% by weight, preferably 0.3% by weight and most preferably 0.1% by weight, since it adversely affects the oxidation behavior of the alloy. This segregation can increase spalling of the protective oxide scale. If the concentration of these unavoidable impurities exceeds a certain level, problems surrounding product yields will arise, as well as degradation of the material properties of the alloy.

[0122] Iron behaves similarly to nickel and can be added as a low-cost alternative to nickel. Furthermore, allowing the addition of iron improves the ability of the alloy to be produced from recycled materials. Therefore, iron is preferably present in an amount of at least 0.1% by weight. However, to significantly reduce costs, iron can be added up to 5.0% by weight. Preferably, the iron addition is no more than 3.8% or even 2.0% by weight to reduce the tendency to form undesirable Laves phases (which reduce the alloy's mechanical properties) and to minimize the very detrimental effect that iron has on oxidation resistance. Most preferably, the iron addition is limited to 1% by weight. This produces an alloy with good ability to be recycled with little loss of material performance.

[0123] Addition of hafnium (Hf) up to 1.0 wt. % is beneficial to bind unavoidable impurities in the alloy and to provide strength. Hafnium is a strong carbide former and may provide additional grain boundary strengthening. However, hafnium has a high elemental cost and may exacerbate solidification cracking by strongly segregating into the final liquid during additive manufacturing. Therefore, it is more preferable to limit the amount of hafnium to 0.5 wt. %, and even more preferable to limit it to 0.25 wt. %.

[0124] The so-called "reactive elements" (yttrium (Y), lanthanum (La) and cerium (Ce)) are added at levels up to 0.1% by mass. 2 O 3These reactive elements can "sweep" harmful elements such as sulfur, which segregates to the alloy-oxide interface and weakens the bond between the oxide and the substrate, leading to oxide spalling. Magnesium (Mg) also exhibits "swabbing" behavior for harmful elements and can have beneficial effects on mechanical properties, so it can be added up to 0.1 wt.%. Silicon (Si) can be beneficially added up to 0.5 wt.%. The addition of silicon to nickel-base superalloys at levels up to 0.5 wt.% has been shown to be beneficial for oxidation properties. Silicon in particular segregates to the alloy / oxide interface and improves the bond of the oxide to the substrate. This reduces oxide spalling, resulting in improved oxidation resistance. (Example of the present invention)

[0125] [Table 5] [Table 6]

[0126] A new alloy example, ABD-900AM, shown in Tables 5 and 6, was designed to overcome the shortcomings of conventional alloys in additive manufacturing. Table 5 shows the nominal composition, in mass percent, of the newly designed high gamma-volume fraction nickel-base superalloy compared to the alloys listed in Table 1. Table 6 shows the phase fractions and figures of merit calculated by the "Alloy Design" software. This is the result of the conventional nickel-base superalloys listed in Table 1. The gamma-content is controlled by varying the value of f(SAC), which is lower than many benchmark alloys, thus reducing the risk of strain age cracking. The strength figure of merit is designed to be equivalent to conventional alloys despite the low gamma-content, due to the high ratio of Nb and Ta, which act to increase the APB strengthening factor. These additions also significantly reduce the hot tearing index, thus reducing the risk of cracking during solidification. The creep figure of merit is preferably maximized within the range of alloy stability by the presence of Mo and W, and favorable oxidation resistance is preferably achieved by a high Cr / Ti ratio.

[0127] Figure 16 shows that the microcrack density is reduced in ABD-900AM compared to Haynes 282 and IN 738. Testing has shown that the strength and creep resistance of ABD-900AM is equivalent to additively manufactured IN 738.

[0128] [Table 7] [Table 8]

[0129] Table 7 shows the nominal composition in mass% of the newly designed nickel-base superalloys with high gamma-proportion volume fraction. The levels of C and B are fixed at 0.04% and 0.005% by mass, respectively. Table 8 shows the phase percentages and figures of merit in the alloys listed in Table 7, calculated with the "Alloy Design" software. Tables 7 and 8 show variations of ABD-900AM with modified Fe and Co contents. AM901 and AM902 contain small amounts of Fe (used as a substitute for Co), which increases the resistance to recycled materials and slightly reduces element costs, at the expense of some oxidation resistance. In AM903 to AM908, the content of Co (used as a substitute for Ni) is gradually decreased. This has the beneficial effect of reducing the cost of the alloys, since the price of Co is relatively high. However, the lower cobalt levels slightly increase the hot tearing index and slightly decrease the gamma-proportion content. AM 906 to AM 908 have the lowest strength figures of merit among all the examples. AM 906 through AM-908 have low levels of cobalt, an element that improves strength, as well as intermediate temperature creep resistance and thermal conductivity.

[0130] [Table 9] [Table 10]

[0131] Table 9 shows the nominal composition in wt.% of the newly designed high gamma prime volume fraction nickel-base superalloys. The C and B levels are fixed at 0.04 wt.% and 0.005 wt.%, respectively. Table 10 shows the phase fractions and merit figures for the alloys listed in Table 9, calculated with the "Alloy Design" software. Tables 9 and 10 show variations of ABD-900AM optimized to enhance grain boundary strength with the addition of Zr and Hf. AM911 and AM912 may also have excellent oxidation resistance due to the beneficial effect of Hf on oxide scale adhesion. However, all of these alloys have higher hot tearing index and wider shale solidification range than the base ABD-900AM composition due to the action of Zr and Hf as melting point depressants. Thus, there is a trade-off to be managed between the printability and grain boundary strength of the alloys within the scope of the present invention.

[0132] [Table 11] [Table 12]

[0133] Table 11 shows the nominal composition in wt.% of the newly designed high gamma prime volume fraction nickel-base superalloys. The C and B levels are fixed at 0.04 wt.% and 0.005 wt.%, respectively. Table 12 shows the phase percentages and figures of merit for the alloys listed in Table 11, as calculated by the "Alloy Design" software. Tables 11 and 12 show alloys AM913 to AM917, which are variations of ABD-900 with the addition of Re and Ru to enhance creep resistance. Of the two elements, Re is seen to have a greater impact in terms of creep figure of merit while at the same time having a less adverse effect on cost and hot tearing index. However, due to the high atomic weight of Re, the addition of Re significantly increases density.

[0134] [Table 13] [Table 14]

[0135] Table 13 shows the nominal composition in weight percent of the newly designed high gamma prime volume fraction nickel-base superalloys. The levels of C and B are fixed at 0.04% and 0.005% by weight, respectively. Table 14 shows the phase percentages and figures of merit for the alloys listed in Table 13, as calculated by the "Alloy Design" software. Tables 13 and 14 show a series of variations of ABD-900AM in which the amounts and ratios of the gamma prime phase formers have been manipulated. Alloys AM918 and AM919 contain increasing levels of niobium and tantalum. Niobium and tantalum are used as substitutes for titanium on a density-corrected basis to maintain a constant strain age cracking figure of merit. This results in a gradual decrease in the hot cracking index and an increase in the strength figure of merit and oxidation resistance (as expressed by the Cr / Ti ratio). AM921 and AM922 have proportional increases in the amounts of Al, Nb, Ti, and Ta (on a density-corrected basis). The result is alloys with high gamma prime content and strength figures of merit, as well as high strain age cracking figures of merit. These alloys are expected to have low oxidation resistance in the preferred range due to the low Cr / Ti ratio.

[0136] [Table 15] [Table 16]

[0137] Table 15 shows the nominal composition in wt.% of the newly designed high gamma-volume fraction nickel-base superalloys. The levels of C and B are fixed at 0.04 wt.% and 0.005 wt.%, respectively. Table 16 shows the phase proportions and figures of merit for the alloys listed in Table 15, calculated with the "Alloy Design" software. Tables 15 and 16 show a series of alloys in which the Cr / Ti ratio has been manipulated. Alloys AM922-AM925 have the same ratio as ABD-900AM, but with proportionally less amounts of both Cr and Ti. This may improve oxidation performance and improve mechanical properties. Alloys AM922 and AM923 keep the gamma-content nearly constant by slightly increasing the percentages of other gamma-formers, while increasing the levels of Mo and W to maintain the same Md number. Alloys AM924 and AM925 increase the gamma-content even further than that of ABD-900AM, while the Mo and W contents remain unchanged.

[0138] Alloy AM926 is a modification of ABD-900AM, but with a reduced titanium content to achieve a Cr / Ti ratio of 10 for improved oxidation resistance, while proportionately increasing the levels of other gamma-prime forming elements to maintain the same gamma prime fraction. Alloys AM927-AM930 are homologous to AM922-AM927, but with a fixed Cr / Ti ratio of 10. Similarly, alloys AM931-AM933 are homologous to AM926-AM930, but with a Cr / Ti ratio of 14. In alloys of this ratio, a Cr content of 13 is not considered, as this falls below the Ti level required for the present invention.

[0139] [Table 17] [Table 18]

[0140] Table 17 shows the nominal composition in wt.% of the newly designed high gamma prime volume fraction nickel-base superalloys. The C and B levels are fixed at 0.04% and 0.005% by weight, respectively. Table 18 shows the phase fractions and figures of merit for the alloys listed in Table 17, as calculated by the "Alloy Design" software. Tables 17 and 18 show a series of alloys (AM934-AM945) directly equivalent to Tables 15 and 16, with increasing gamma prime content by proportionally increasing the function f(SAC). This shows a range of chromia forming alloys suitable for additive manufacturing, with oxidation resistance, high gamma prime and low hot tearing index.

Claims

1. 1.5 to 4.5% by mass of aluminum, 1.1 to 3.4% by mass of titanium, 0.0 to 4.0% by mass of niobium, 0.0 to 5.2% by mass of tantalum, 0.9 to 6.6% by mass of tungsten, 0.0 to 3.0% by mass of molybdenum, 0.0 to 24.0% by mass of cobalt, 12.5 to 20.6% by mass of chromium, 0.02 to 0.15% by mass of carbon, 0.001 to 0.015% by mass of boron, 0.0 to 0.1% by mass of zirconium, 0.0 to 3.0% by mass of rhenium, 0.0 to 2.0% by mass of ruthenium, 0.0 to 3.0% by mass of iridium, 0.0 to 0.5% by mass of vanadium, 0.0 to The alloy contains 1.0 mass% palladium, 0.0 to 1.0 mass% platinum, 0.0 to 0.5 mass% silicon, 0.0 to 0.1 mass% yttrium, 0.0 to 0.1 mass% lanthanum, 0.0 to 0.1 mass% cerium, 0.0 to 0.003 mass% sulfur, 0.0 to 0.25 mass% manganese, 0.0 to 0.1 mass% magnesium, 0.0 to 0.5 mass% copper, 0.0 to 1.0 mass% hafnium, and the balance is nickel and unavoidable impurities. Iron is not contained, and the mass percentages of aluminum, titanium, niobium, tantalum, tungsten, and molybdenum contained in the alloy are respectively W. Ａｌ , W Ｔｉ , W Ｎｂ , W Ｔａ , W Ｗ and W Mo, satisfying the following formula: 0.65≦0.3W Ｎｂ +0.15W Ｔａ 3.6≦W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ ≦5.7 W Ｔａ +0.92W Ｗ ≦6.1 W W +0.77W Mo ≧3.0

2. 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, which satisfies the following formula: 4.1≦W Ａｌ +0.5W Ｔｉ +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: W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ ≦5.5

4. 4. The nickel-base alloy composition of claim 1, comprising a volume fraction of gamma prime phase of 44% or less at 900°C.

5. 5. The nickel-base alloy composition of claim 1, comprising a volume fraction of gamma prime phase of 19% or greater at 900°C.

6. The nickel-based alloy composition according to claim 1 , comprising, by mass%, 16.0% or more of chromium.

7. 7. The nickel-based alloy composition of claim 1, comprising, by weight, up to 19.5% chromium.

8. 8. The nickel-based alloy composition of claim 1, comprising, by weight, tantalum in an amount up to 4.2%.

9. 9. The nickel-based alloy composition of claim 1 , comprising, by weight, up to 2.9% molybdenum.

10. 10. The nickel-based alloy composition of claim 1, comprising, by weight percent, titanium in an amount up to 3.0%.

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

12. 12. The nickel-based alloy composition of claim 1, comprising up to 5.5% tungsten, by weight.

13. 13. The nickel-based alloy composition of claim 1, comprising, by weight, niobium in an amount up to 3.0%.

14. 14. The nickel-based alloy composition of claim 1, further comprising, by weight, up to 0.5% of one or both of platinum and palladium.

15. 15. The nickel-based alloy composition according to claim 1, comprising, by weight percent, 1.7% or more of aluminum.

16. 16. The nickel-based alloy composition of claim 1 , comprising, by weight percent, up to 3.9% aluminum.

17. 17. The nickel-based alloy composition of claim 1, comprising, by weight, up to 20.0% cobalt.

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

19. 19. The nickel-based alloy composition according to claim 1 , comprising, by weight percent, tantalum in an amount of at least 0.6%.

20. 20. The nickel-based alloy composition according to claim 1, comprising, by weight percent, tungsten in an amount of 1.2% or more.

21. 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 20, which satisfies the following formula: W Ａｌ +0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ +0.125W Ｐｔ +0.225W Ｐｄ ≦5.7

22. 22. The nickel-based alloy composition of claim 1, comprising, by weight, at least 5.0% cobalt.

23. 23. The nickel-based alloy composition according to claim 1, comprising, by weight, niobium in an amount of at least 0.6%.

24. The mass percentages of tungsten, molybdenum and chromium are W Ｗ , W Ｍｏ and W Cｒ The nickel-based alloy composition according to any one of claims 1 to 23, which satisfies the following formula: <h2 style=";text-align:left;direction:ltr">W<h2 style=";text-align:left;direction:ltr"> Ｗ <h2 style=";text-align:left;direction:ltr"> +0.74W<h2 style=";text-align:left;direction:ltr"> Mo <h2 style=";text-align:left;direction:ltr"> 2020<h2 style=";text-align:left;direction:ltr"> Cｒ <h2 style=";text-align:left;direction:ltr"> ≦2211

25. The mass percentages of niobium, tantalum, titanium and aluminum are W Ｎｂ , W Ｔａ , W Ｔｉ and W Ａｌ The nickel-based alloy composition according to any one of claims 1 to 24, which satisfies the following formula: (0.5W Ｔｉ +0.3W Ｎｂ +0.15W Ｔａ ) / W Ａｌ ≦1.5

26. The mass percentages of chromium and titanium are W Ｃｒ and W Ｔｉ The nickel-based alloy composition according to any one of claims 1 to 25, which satisfies the following formula: W Ｃｒ / W Ｔｉ ≧6.0

27. The mass percentages of tantalum and tungsten are W Ｔａ and W Ｗ The nickel-based alloy composition according to any one of claims 1 to 26, which satisfies the following formula: W Ｔａ +0.92W Ｗ ≦5.4

28. The mass percentages of niobium and tantalum are W Ｎｂ and W Ｔａ The nickel-based alloy composition according to any one of claims 1 to 27, which satisfies the following formula: 0.68≦0.3W Ｎｂ +0.15W Ｔａ

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