JP2022542384A - Nickel-based alloy - Google Patents

Abstract

17.0-21.3% Chromium, 7.1% Cobalt, 0.9-6.3% Molybdenum, 4.9% Tungsten, 1.8-3.2%, by weight aluminum, 1.8-4.0% titanium, ≤3.05% tantalum, ≤3.0% niobium, ≤0.1% carbon, 0.001-0.1% boron, 0 .001-0.5% zirconium, 0.02% or less magnesium, 0.5% or less silicon, 0.1% or less yttrium, 0.1% or less lanthanum, 0.1% or less cerium, 0.003% or less sulfur, 0.25% or less manganese, 0.5% or less copper, 0.5% or less hafnium, 0.5% or less vanadium, and 10.0% or less iron, A nickel-based alloy composition, the balance being nickel and incidental impurities.

Description

The present invention relates to cast and forged (C&W) nickel-base superalloy compositions for use in high stress high temperature applications above 650° C., such as rotating components in gas turbine engines and other turbomachinery.

Improving alloy performance in terms of maximum operating temperature and maximum service life can have a significant impact on engine efficiency and the cost effectiveness of engine operation.

Examples of typical compositions of C&W nickel-base superalloys used in high temperature and high stress applications are listed in Table 1. The development of alloys with higher strength has tended to reduce alloy workability and increase the use of higher cost elements. Table 1 is the nominal composition in weight percent of cast and forged gamma/gamma prime nickel-based superalloys.

An object of the present invention is to achieve an improved trade-off between mechanical strength and hot workability. In certain cases, it is desired to improve the mechanical strength compared to alloys such as Waspaloy and Haynes 282, and the strength of high strength alloys such as Rene65, AD730, 720Li, TMW-4 is targeted. This was achieved by increasing the γ' volume fraction to higher levels than Waspaloy or Haynes282, as explained in connection with FIG. In certain cases, lower γ' solvus than the current high strength alloys Rene 65, AD730, 720Li and TMW-4 is desired to maintain better hot workability. In its optimum form, the alloy of the present invention simultaneously achieves both improved mechanical strength compared to Waspaloy and Haynes 282 and lower solvus than Rene65, AD730, 720Li and TMW-4.

According to the invention, in % by weight, 17.0-21.3% chromium, 7.1% or less cobalt, 0.9-6.3% molybdenum, 4.9% or less tungsten,1. 8-3.2% aluminum, 1.8-4.0% titanium, 3.05% or less tantalum, 3.0% or less niobium, 0.1% or less carbon, 0.001-0. 0.001-0.5% zirconium; 0.02% or less magnesium; 0.5% or less silicon; 0.1% or less yttrium; 0.1% or less lanthanum; not more than 1% cerium, not more than 0.003% sulfur, not more than 0.25% manganese, not more than 0.5% copper, not more than 0.5% hafnium, not more than 0.5% vanadium and 10.0% A nickel-based alloy composition is provided comprising the following iron, with the balance being nickel and incidental impurities: Such alloys provide an improved trade-off between strength and gamma prime solvus temperature.

In one embodiment, the weight percentages of aluminum, titanium, niobium, and tantalum in the alloy are W _Al , W _Ti , W _Nb , and W _Ta , respectively, satisfying the following equations.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≤ 1.1
Preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≤ 1.0
Such alloys are less likely to form eta and delta phases, especially when produced by casting and forging processes.

In one embodiment, the weight percentages of aluminum, titanium, niobium, and tantalum in the alloy are W _Al , W _Ti , W _Nb , and W _Ta , respectively, satisfying the following equations.
(0.6W _Ti + 0.31W _Nb + 0.15W _Ta )/ _WAl ≥ 0.64
Preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.72
More preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.76
Even more preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.93
Such alloys have improved mechanical strength.

In one embodiment, the volume fraction of gamma prime phase in the alloy is 48% or less at 760°C, preferably 44% or less at 760°C, more preferably 40% or less at 760°C, Most preferably, it is 36% or less at 760°C. This is effective in limiting the gamma prime solvus, thus improving hot workability.

In one embodiment, the weight percentages of aluminum, titanium, tantalum, and niobium in the alloy are W _Al , W _Ti , W _Ta , and W _Nb , respectively, satisfying the following equations.
3.4 ≤ 0.6W _Ti + 0.31W _Nb + 0.15W _Ta + 0.79W _Al
Preferably, the following formula is satisfied.
3.6 ≤ 0.6W _Ti + 0.31W _Nb + 0.15W _Ta + 0.79W _Al
Such alloys have acceptable levels of gamma prime and are therefore strong.

In one embodiment, the weight percentages of aluminum, titanium, tantalum, and niobium in the alloy are W _Al , W _Ti , W _Ta , and W _Nb , respectively, satisfying the following equations.
4.6≧0.6W _Ti +0.31W _Nb +0.15W _Ta +0.79W _Al
Preferably, the following formula is satisfied.
4.3≧0.6W _Ti +0.31W _Nb +0.15W _Ta +0.79W _Al
More preferably, the following formula is satisfied.
4.0≧0.6W _Ti +0.31W _Nb +0.15W _Ta +0.79W _Al
Such alloys have better workability.

In one embodiment, the nickel-base alloy composition comprises 18.0% or more by weight of chromium. Such alloys have improved resistance to oxidation and corrosion.

In one embodiment, the nickel-based alloy composition comprises 19.5% or less by weight of chromium. Such alloys are less likely to precipitate harmful phases.

In one embodiment, the nickel-based alloy composition comprises, by weight percent, tantalum of 2.25% or less, preferably 1.5% or less, and most preferably 0.65% or less. Such alloys can have improved mechanical strength and/or lower levels of niobium for the same strength.

In one embodiment, the molybdenum content of the nickel-base alloy composition is 2.0% or more, preferably 3.0% or more, by weight. Such alloys provide an improved combination of increased creep resistance and low density.

In one embodiment, the nickel-based alloy composition comprises, by mass %, titanium of 1.9% or more, preferably 2.0% or more, more preferably 2.3% or more, and even more preferably 2.4%. above, and most preferably above 2.5%. Such alloys have a low gamma prime solvus temperature.

In one embodiment, the nickel-based alloy composition comprises, by weight percent, tungsten of 2.7% or less, preferably 0.7% or less. Such alloys have low densities.

In one embodiment, the nickel-base alloy composition comprises 2.0% or less niobium by weight. Such alloys have better oxidation resistance.

In one embodiment, the nickel-based alloy composition comprises, by weight percent, titanium of 3.5% or less, preferably 3.0% or less. Such alloys have better oxidation resistance.

In one embodiment, the nickel-based alloy composition comprises 1.9% or more by weight of aluminum. Such alloys have improved stability of the gamma prime phase.

In one embodiment, the nickel-based alloy composition comprises, by weight percent, aluminum of 2.75% or less, preferably 2.6% or less, more preferably 2.3% or less. Such alloys provide a better combination of strength and hot workability.

In one embodiment, the nickel-based alloy composition comprises, by weight percent, cobalt of 5.3% or less, preferably 3.5% or less, and more preferably 1.6% or less. Such alloys are less costly.

In one embodiment, the mass percentages of tantalum and cobalt contained in the alloy are W _Ta and W _Co , respectively, and satisfy the following equations.
W _Ta +0.43 W _Co ≦3.05
Preferably, the following formula is satisfied.
W _Ta +0.43 W _Co ≦2.25
More preferably, the following formula is satisfied.
W _Ta +0.43 W _Co ≦1.5
Most preferably, the following formula is satisfied.
W _Ta +0.43 W _Co ≦0.65
Such alloys limit the cost.

In one embodiment, the weight percentages of titanium, niobium, tantalum, and aluminum in the alloy are W _Ti , W _Nb , W _Ta , and W _Al , respectively, satisfying the following equations.
0.6W _Ti + 0.44 ( _WNb + 0.66W _Ta ) + 0.2W _Al ≥ 2.8
Such alloys have better mechanical strength.

In one embodiment, the nickel-base alloy composition comprises, by weight percent, niobium of 0.3% or more, preferably 0.4% or more, more preferably 0.7% or more, and even more preferably 0.8%. above, and most preferably above 1.2%. Such alloys have superior strength.

In one embodiment, the weight percentages of molybdenum and tungsten contained in the alloy are W _Mo and W _W , respectively, and satisfy the following equations.
W _Mo +0.5 _{W W} ≧3.4
Preferably, the following formula is satisfied.
W _Mo +0.5 _{W W} ≧3.6
Such alloys have superior high temperature properties such as tensile strength and creep resistance due to the stronger matrix due to solution hardening.

In one embodiment, the weight percentages of molybdenum and tungsten contained in the alloy are W _Mo and W _W , respectively, and satisfy the following equations.
_WMo + _0.5WW ≦6.3
Preferably, the following formula is satisfied.
_WMo + _0.5WW≤5.6
More preferably, the following formula is satisfied.
_WMo + _0.5WW≤5.1
Most preferably, the following formula is satisfied.
_WMo + _0.5WW≤4.3
Such alloys have good high temperature creep strength due to the strong matrix phase.

In one embodiment, the volume fraction of gamma prime phase in the alloy is at least 29% at 760°C, preferably at least 31% at 760°C, and most preferably at least 40% at 760°C. Such alloys have good tensile strength properties.

The term "comprising" is used herein to indicate that the composition is 100% and that the presence of additional ingredients is excluded to bring the percentage to 100%. Unless stated otherwise, % is expressed as % by weight.

The invention will now be more fully described, by way of example only, with reference to the accompanying drawings.

FIG. 1 shows the calculated correlation between the γ′ volume fraction at 760° C. and the γ′ solvus temperature for alloys within the alloy design region listed in Table 2. The dashed line indicates the preferred limit for the γ' solvus temperature. FIG. 2 shows the calculated correlation between the volume fraction of the γ′ phase at 760° C. and the alloy strength (in terms of strength figure of merit). The calculated strength of high-strength alloys such as Rene65, AD730, 720Li and TMW-4 are shown in dashed lines. FIG. 3 shows the relationship between alloy strength and elemental proportions based on the formula ((0.6W _Ti +0.31W _Nb +0.15W _Ta )/W _Al ). The dashed lines plotted on the graph indicate the intensity limit for various preferred ranges of γ' volume fraction. The calculated strength of high-strength alloys such as Rene65, AD730, 720Li and TMW-4 are shown in dashed lines. FIG. 4 shows the sum of the elemental aluminum and the elements niobium, tantalum and titanium (based on the relationship 0.6W _Ti + 0.31W _Nb + 0.15W _Ta ) versus the γ' volume fraction at 760°C; is a contour map showing the influence of . Superimposed on FIG. 4 are preferred limits for the proportions of the elements titanium, tantalum, niobium and aluminum based on the relationship (0.6 W _Ti +0.31 W _Nb +0.15 W _Ta )/W _Al . The shaded area defines the target composition range in this application. FIG. 5 is a contour plot showing the effect of the elements aluminum and titanium on yield strength (in terms of strength figure of merit). In this alloy, the niobium content is fixed at 0.0% by weight and the tantalum content is fixed at 0.0% by weight. Superimposed on FIG. 5 are the preferred limits for the sum of the titanium, tantalum, niobium and aluminum elements based on their relationship to 0.79 W _Al +0.6 W _Ti +0.31 W _Nb +0.15 W _Ta , which indicates various levels of γ′ volume fractions. Also superimposed on FIG. 5 are the preferred limits of (0.6 W _Ti +0.31 W _Nb +0.15 W _Ta )/W _Al . The shaded area defines the target composition range in this application. FIG. 6 is a contour plot showing the effect of the elements aluminum and titanium on yield strength (in terms of strength figure of merit). In this alloy, the niobium content is fixed at 1.0% by weight and the tantalum content is fixed at 0.0% by weight. Superimposed on FIG. 6 are the preferred limits for the sum of the elements titanium, tantalum, niobium and aluminum based on their relationship to 0.79 W _Al +0.6 W _Ti +0.31 W _Nb +0.15 W _Ta , which indicates various levels of γ′ volume fractions. Also superimposed on FIG. 6 is the preferred limit of (0.6 W _Ti +0.31 W _Nb +0.15 W _Ta )/W _Al . The shaded area defines the target composition range in this application. FIG. 7 is a contour plot showing the effect of the elements aluminum and titanium on yield strength (in terms of strength figure of merit). In this alloy, the niobium content is fixed at 2.0% by weight and the tantalum content is fixed at 0.0% by weight. Superimposed on FIG. 7 are the preferred limits for the sum of the titanium, tantalum, niobium and aluminum elements based on their relationship to 0.79 W _Al +0.6 W _Ti +0.31 W _Nb +0.15 W _Ta , which indicates various levels of γ′ volume fractions. Also superimposed on FIG. 7 are the preferred limits of (0.6 W _Ti +0.31 W _Nb +0.15 W _Ta )/W _Al . The shaded area defines the target composition range in this application. FIG. 8 is a contour plot showing the effect of the elements aluminum and titanium on yield strength (in terms of strength figure of merit). In this alloy, the niobium content is fixed at 3.0% by weight and the tantalum content is fixed at 0.0% by weight. Superimposed on FIG. 8 are the preferred limits for the sum of the titanium, tantalum, niobium and aluminum elements based on their relationship to 0.79 W _Al +0.6 W _Ti +0.31 W _Nb +0.15 W _Ta , which indicates various levels of γ′ volume fractions. Also superimposed on FIG. 8 is the preferred limit of (0.6 W _Ti +0.31 W _Nb +0.15 W _Ta )/W _Al . The shaded area defines the target composition range in this application. FIG. 9 is a contour plot showing the effect of elemental molybdenum and elemental tungsten (based on the relationship W _Mo +0.5 _{W W} ) on the solid solubility index. FIG. 10 shows the sum of the γ' volume fraction and the sum of the molybdenum and tungsten elements (based on the relationship between W _Mo +0.5 _{W W} ) for the maximum chromium content that can achieve the target stability number (0.90). and the influence of FIG. 11 shows the sum of the γ′ volume fraction and the sum of the elements molybdenum and tungsten (based on the relation W _Mo +0.5 _{W W} ) for the maximum chromium content capable of achieving the target stability number (0.89). and the influence of FIG. 12 is a contour plot showing the effect of cobalt and tantalum elements on alloy cost. FIG. 13 is a contour plot showing the effect of tungsten and tantalum elements on alloy density.

Traditionally, nickel-base superalloys have been designed on an empirical basis. Accordingly, the chemical composition of nickel-base superalloys has been identified by 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 employed. The large number of alloying elements that can achieve this combination indicates that these alloys have not been fully optimized and that there are likely to be improved alloys.

Chromium (Cr) and aluminum (Al) are typically added in superalloys to provide oxidation/corrosion resistance, and cobalt (Co) is added to improve resistance to sulfidation. Molybdenum (Mo), tungsten (W) and cobalt (Co) are introduced for creep resistance, since these elements undergo thermal activation processes (e.g. dislocations) that determine the rate of creep deformation. increase). Aluminum (Al), tantalum (Ta), niobium (Nb) and titanium (Ti) are introduced to increase static and cyclic strength, since these elements form the precipitation hardening phase gamma prime ( γ') is to be promoted. This precipitated phase is coherent with a face centered cubic (FCC) matrix phase called gamma (γ).

The model-based approach used to identify new grades of nickel-base superalloys is described herein under the term "alloy design" (ABD) method. This approach utilizes a computational materials modeling framework for estimating design-related properties over a very wide range of compositions. In principle, this alloy design tool enables the so-called inverse problem to be solved. That is, an optimum alloy composition can be identified that best satisfies specified design constraints.

The first step in the design process is to define a table of elements and upper and lower compositional limits associated with the table of elements. In the present invention, the composition limit for each element when adding each element, called "alloy design area", is taken into consideration. This compositional limit is detailed in Table 2. Table 2 shows the alloy design regions that were studied.

The second step is based on thermodynamic calculations to calculate the phase diagram and thermodynamic properties of a particular alloy composition. This is often called the CALPHAD method (CALculate PHAse Diagram). Performing these calculations at the typical operating temperature of the new alloy (900° C.) provides information about the phase equilibrium (microstructure).

The third step involves identifying the alloy composition with the desired microstructure. For nickel-based superalloys that require superior resistance to creep deformation, creep rupture life is progressively improved as the volume fraction of the precipitation hardening phase γ' increases. The range of γ′ volume fractions for which creep rupture life is most beneficial is 60-70% at 900° C. (although other design constraints limit the volume fraction to lower values). alloys with a γ′ volume fraction of 50% to 60% are included). When the volume fraction of γ' exceeds 70%, a decrease in creep resistance is observed.

Also, the γ/γ′ lattice mismatch loses coherency and must follow positive or negative values, whichever is smaller. Therefore the limit depends on the absolute value of that value. The lattice mismatch δ is defined as the mismatch between the γ phase and the γ' phase and is obtained by the following formula.

where α _γ and α _γ ' are the lattice constants of the γ and γ' phases.

Therefore, the model identifies all compositions within the design space for which the calculated volume fraction of γ' results in the desired value. In these compositions, the lattice mismatch of γ' is less than a certain absolute value.

In a fourth step, figures of merit are estimated for the identified alloy compositions remaining in the dataset. The indices of merit include creep merit index (indicating the creep resistance of an alloy based on average composition only), strength merit index (indicating an alloy's precipitation yield strength based on average composition only), density, Included are cost, microstructural stability and gamma prime solvus temperature.

In the fifth stage, the calculated figure of merit is compared to the constraints on the desired behavior and these design constraints are taken 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 dataset is very small.

The sixth and final step involves analyzing the remaining composition data set. This analysis can be done in various ways. On the one hand, alloys with the highest figure of merit may be sorted through a database. Alloys with the highest figure of merit are, for example, the lightest alloys, the most creep resistant alloys, the most oxidation resistant alloys, and the least expensive alloys. Alternatively, a database may be used to determine the relative performance trade-offs made by different combinations of properties.

Explain the seven merit indices.

The first figure of merit is the creep figure of merit. The most important observation is that the time-dependent deformation (ie, creep) of nickel-based superalloys is caused by dislocation creep associated with initial activity confined to the γ phase. Therefore, the dislocation segment is rapidly anchored at the γ/γ' interface due to the increased proportion of the γ' phase. The rate-limiting step is the departure 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) which causes the critical effect of the alloy composition on the creep properties.

A physics-based microstructural model is employed for the rate of creep strain ε ^· accumulation when the loading is uniaxial and along the <001> crystallographic direction. The set equation is the following formula.

where 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 a constraint factor.

The term K _CF indicates the proximity of the cubic grains within these alloys. Equation 3 describes the dislocation multiplication process that requires an estimate of the multiplication parameter C and the initial dislocation density. The term D _eff is the effective diffusivity that controls the rising process at the particle/matrix interface.

Note that in the above content the composition dependence arises from the two terms φ _p and D _eff . Therefore, assuming a constant microstructure (much of which is controlled by heat treatment), a dependence on chemical composition is caused by D _eff since φ _p is fixed. It will be appreciated that for the purposes of alloy design modeling as described herein, it is not necessary to perform the full integration of Equations 2 and 3 for each prototype alloy composition. Instead, a first-order figure of merit M _creep is used, which needs to be maximized. M _creep is obtained by the following formula.

where x _i is the atomic fraction of solute i in the γ phase. D _i ^~ is the appropriate interdiffusion coefficient.

The second figure of merit is the strength figure of merit. For high nickel-base superalloys, most of the strength comes from the precipitation phases. Therefore, optimizing the alloy composition for maximum precipitation strength is an important design consideration. Based on hardening theory, a figure of merit for strength, 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 formula:

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.

From equation (5), it is clear that the defect energy in the γ' phase (eg, the antiphase boundary APB energy) has a large impact on the deformation behavior of nickel-based superalloys. Increasing the APB energy was found to improve mechanical properties, including tensile strength and resistance to creep deformation. APB energy studies have been performed for many Ni--Al--X systems using density functional theory. This study calculated the effect of ternary elements on the APB energy of the γ′ phase and assumed a linear superposition of the effects of each ternary addition when considering complex multicomponent systems. As a result, the following formula was derived.

where x _Cr , x _Mo , x _W , x _Ta , x _Nb and x _Ti respectively represent the atomic % concentrations of chromium, molybdenum, tungsten, tantalum, niobium and titanium in the γ′ phase. The composition in the γ' phase is determined by phase equilibrium calculation.

A third figure of merit is density. The density ρ was calculated using a simple rule for mixtures and a correction factor. where ρ _i is the density of the given element and x _i is the atomic fraction of the alloying element.

A fourth figure of merit is cost. A simple rule of mixtures was applied to estimate the cost of each alloy. Here, the cost of each alloy is obtained by multiplying the mass fraction x _i of the alloying element by the current (2017) raw material cost c _i of the alloying element.

This estimate assumes that processing costs are the same for all alloys. Thus, product yield is not affected by composition.

A fifth figure of merit is based on the exclusion of alloy candidates based on inappropriate microstructures created based on their susceptibility to TCP phases. To do this, the d-orbital energy levels (referred to as Md) of the alloying elements are used to determine the total effective Md level according to the following equation.

where x _i represents the mole fraction of element i contained in the alloy. A higher value of Md indicates a higher probability of TCP formation.

The sixth figure of merit is the gamma prime solvus temperature. The gamma prime solvus is defined as the temperature at which the gamma prime volume fraction tends to zero. This is determined using thermodynamic calculations as described above in the second step of the alloy design method. It is used to calculate the phase diagram and thermodynamic properties of a particular alloy composition and to find the temperature at which phase transitions occur.

The seventh figure of merit is the solid solution figure of merit. Solid-solution hardening occurs in a (FCC) matrix phase called gamma (γ), and especially this hardening mechanism is important at high temperatures for high strength and creep resistance. A model is adopted that assumes the superposition of individual solute atoms on matrix phase strengthening. The solid solution strength coefficients k _i of the elements considered within the design area are 225, 39.4, 337, 1015, 1183, 1191, 775 for aluminum, cobalt, chromium, molybdenum, niobium, tantalum, titanium and tungsten, respectively. and 977 MPa/at.% ^1/2 (H. Roth, C. Davis, and R. Thomson: Metallurgical and Materials Transactions A, 1997, vol. 28, pp. 1329-1335). The solid solubility index is calculated based on the equilibrium composition of the matrix phase using the following formula.

Here, M _{solid-solution} is the solid-solution merit index, and x _i is the concentration of element i in the γ matrix phase.

The alloy composition of the present invention was identified using the ABD method described above. The design intent of this alloy is to develop a cast and forged (C&W) type alloy with an improved balance of high strength and good hot workability. This is combined with reduced alloy costs, high levels of oxidation and corrosion resistance, excellent microstructural stability, and excellent alloy densities. The new alloy's balance of properties provides significant advantages over other C&W alloys described in the prior art, especially for use in structural applications where stresses are high and temperatures above 650°C.

Material properties (determined using the ABD method) are listed in Table 3 at nominal compositions (listed in Table 1) for commercial C&W alloys used at high stress and temperature. New alloy designs were considered in conjunction with the predicted properties listed for these alloys. Table 3 shows the calculated phase fractions, mismatches and figures of merit produced by the "alloy design" software. This is the result for the nickel-based superalloys listed in Table 1.

The design principles of the new alloy are described below.

In nickel-base superalloys used in cast and forged (C&W) forms, there is a balance between hot workability ("hot workability") and strength of the alloy. FIG. 1 shows that, in general, there is a correlation between increasing γ', the major strengthening phase, and increasing γ' solvus temperature. Typically, the alloys of the present invention are hot worked at temperatures near or above the solvus temperature. Therefore, the higher the γ' solvus, the more difficult the hot working is because the material is more resistant to deformation at high temperatures. In addition, increased hot working temperatures can increase thermally induced stresses that build up upon cooling after hot working. This can lead to cracking upon cooling of the hot worked material, such as quench cracking or strain age cracking mechanisms.

In FIG. 2 it can be seen that the increase in strength (determined by the strength merit index) at the desired operating temperature of the alloy (approximately 700-750° C.) correlates with the increase in γ′ solvus temperature. Thus, Figures 1 and 2 show that there is a complex trade-off between the ability of the alloy to be produced in a hot working process and overall strength during operation.

Figures 1 and 2 show alloys within the alloy design space defined in Table 2 that have an improved combination of hot workability and high temperature strength over the current alloys listed in Table 1, albeit with trade-offs. indicates that there is composition. An object of the present invention is to achieve improved mechanical strength compared to alloys such as Waspaloy or Haynes 282, preferably comparable to high strength alloys such as Rene65, AD730, 720Li and TMW-4. Aim for strength. This is achieved by raising the γ' volume fraction to a higher level than Waspaloy or Haynes282, as explained in connection with FIG. A lower solvus than the current high strength alloys (Rene65, AD730, 720Li and TMW-4) combined with improved mechanical strength is desired. This maintains better hot workability.

FIG. 1 shows the correlation between γ′ volume fraction and γ′ solvus temperature in the compositional regions listed in Table 2. It can be seen that for a given γ' volume fraction, there is a spread of γ' solvus temperatures of up to about 100°C. For alloys with a good combination of high temperature strength and hot workability, it is preferred to achieve a maximum value for the γ' volume fraction and a minimum value for the γ' solvus temperature. The solvus for alloys Rene65 and AD730 is 1100°C. In the present invention, it is desirable to have a solvus below 1080°C to ensure improved moldability. Therefore, the γ' volume fraction is limited to 48% at 760°C. In the present invention, it is desirable to limit the γ' volume fraction to 44% or less to ensure a combination of strength, oxidation resistance and alloy microstructural stability. This will be explained later with reference to FIGS. 10-11. In certain cases it is desirable to limit the γ' solvus to 1050°C to further improve hot workability, more preferably to 1025°C to further improve hot workability. Limit to C. Therefore, the γ' volume fraction is preferably limited to 40% or less, more preferably 36% or less. This may come at the expense of reduced strength.

FIG. 2 shows the trade-off between the volume fraction of γ′ phase and alloy strength in terms of the strength figure of merit. It can be seen that there is a general trend of increasing alloy strength as the γ' volume fraction increases. However, for a given γ' volume fraction, it is possible to vary the strength by choosing the optimum alloy chemistry, as will be explained later with reference to Figures 3 and 5-8. Strength levels for several high-strength C&W superalloys are shown in FIG. It is desirable that the alloys of the present invention have strengths at least equivalent to AD730 or Rene65. To achieve this, a minimum γ' volume fraction of 29% at 760°C is required. Preferably, an alloy with strength comparable to alloy 720Li is desired, more preferably an alloy with strength comparable to alloy TMW-4. Therefore, it is more preferred that the minimum value of γ' is at least 31%, and even more preferred to have a γ' volume fraction of 40% or greater.

FIG. 3 shows the relationship between alloy strength and elemental proportions based on the formula ((0.6 W _Ti +0.31 W _Nb +0.15 W _Ta )/W _Al ). The dashed line drawn in the graph indicates the limit of intensity for various ranges of γ′ volume fractions, as described in the previous section. Replacing the γ'-phase aluminum with the elements titanium, niobium and tantalum increases the APB energy and thus promotes strengthening of the alloy (see Equation 6). However, too high a value can lead to the formation of unwanted phases such as eta and delta phases. To avoid precipitation of these unwanted phases in alloys produced via the C&W production route, the ratio of elements based on the formula ((0.6W _Ti + 0.31W _Nb + 0.15W _Ta )/ _WAl ) is , desirably 1.1 or less, more preferably 1.0 or less. This is of particular relevance to the segregation of elements that occurs during ingot manufacture, as elemental segregation can locally enrich some areas of the alloy to a higher proportion. Elemental segregation is kept to a lower level than can be achieved if the alloy is produced by other methods (such as powder metallurgy) by allowing some regions of the alloy to be locally enriched to a higher proportion. be FIG. 3 shows that for γ′ volume fractions of 44% or less, ratios of 0.64 or greater equal the strength merit index of Rene65 and AD730, indicating that this is the minimum desired level of ratio. ing. Fig. 3 shows that when γ' is in the range (29 to 36%) that further improves hot workability, a ratio of 0.72 or more achieves the same strength as Rene65 and AD730, so this is the desirable minimum ratio. It also indicates that it is a level. In the range where γ' is more hot workable (29-40%), a ratio of 0.76 or higher is the desired minimum level of ratio, as strength levels comparable to alloy 720Li are achieved. When γ' is in the most preferable range (29-36%) to further improve hot workability, a ratio of 0.93 or more achieves a strength level equivalent to that of alloy 720Li. This is the minimum desired level of the ratio as it provides the most favorable balance of .

Figure 4 shows the percentage of alloy γ' (as contour lines) and the sum of titanium, tantalum and niobium (based on the relationship 0.6W _Ti + 0.31W _Nb + 0.15W _Ta ) versus the aluminum-based alloy plotted The relationship between composition and The amounts of titanium, niobium and tantalum added are given weighting factors to convert mass % to "aluminum equivalent". This allows direct comparison of elements with widely different densities. For example, aluminum has a density of 2.7 g/cm ³ whereas titanium has a density of 4.5 g/cm ³ so a factor of 0.6 applies (i.e. 2.7/4 .5=0.6). As with titanium, constants are added to convert the additions of elemental niobium (8.57 g/cm ³ ) and elemental tantalum (16.4 g/cm ³ ) to "aluminum equivalents". That is, the correction factors for niobium and tantalum, determined from the densities of those elements relative to aluminum, are 0.31 and 0.15, respectively. From FIG. 4, it can be seen that the γ' volume fraction is determined by the amount of aluminum, titanium, tantalum and niobium added according to the following formula.
f(γ') = 0.6W _Ti + 0.31W _Nb + 0.15W _Ta + 0.79W _Al

Here, f(γ') is a numerical value, and this numerical value must be 3.4 or more in order to make the gamma prime ratio 29% or more. W _Ti , W _Nb , W _Ta and W _Al are the weight percentages of titanium, niobium, tantalum and aluminum contained in the alloy, respectively. In order to set the gamma prime volume fraction to 31% or more, it is preferable to set the numerical range of f(γ') to 3.6. A chromium content of 17.0% by mass or more provides a combination of high strength and high oxidation and corrosion resistance. This is higher than the high-strength alloys listed in Table 3 (AD730, 720Li and TMW-4), resulting in better oxidation relative to these alloys. It is desirable to achieve a high chromium content while maintaining alloy stability (ie, substantially free of TCP phases). Therefore, it is beneficial to keep the gamma prime percentage below 44% (as explained below with reference to FIGS. 10 and 11). In order to make the gamma prime ratio 44% or less, the numerical value of f(γ') must be 4.6 or less. Therefore, the numerical range of f(γ') is preferably 3.6 to 4.6. In order to further improve the hot workability, it is more preferable to set the γ' solvus temperature of the alloy to 1050°C or lower. As a result, γ' is limited to 40% or less, and the numerical value of f(γ') is 4.3 or less. Even more preferably, the γ' solvus temperature of the alloy should be 1025°C or lower. This limits the γ' ratio to 36% or less. To achieve this, the numerical value of f(γ') must be 4.0 or less.

The γ' volume fraction at 760°C is experimentally determined by the following procedure. After prolonged exposure to 760° C. heat, the sample (e.g. 1 cm ³ ) is quenched with water and a portion of the material is cut off using conventional/standard metallurgical preparation techniques using scanning electron microscopy. and polished. When the preparation is completed, the microstructure of γ/γ' is made observable by a scanning electron microscope, and particles with a diameter of 30 nm or more can be observed. A minimum of 10 images are taken, providing a statistically representative data set. The image should cover an area of at least 1 ^mm2 . A two-dimensional image revealing the γ/γ' microstructure must be processed to identify the γ' phase and the area fraction of the gamma prime phase calculated. The area fraction of the phase is taken as the γ' volume fraction.

From FIG. 4, to achieve the combination of strength, workability and gamma prime phase stability (achieved when (0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≤ 1.1) The minimum aluminum content required is determined to be 1.8% by weight or more. Preferably, when (0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≤ 1.0 (to increase the stability of the γ' phase), the minimum aluminum content is 1.9 mass% or more is limited to When the γ′ volume fraction is 44% and (0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.64, the maximum aluminum content in the alloy is up to 3.2 wt%. Limited. Preferably, when γ' is less than or equal to 40% and (0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.76, the aluminum content should be limited to 2.75 wt%. There is

Most preferably, when γ' is 36% or less, (0.6W _Ti + 0.31W _Nb + 0.15W _Ta )/ _WAl ≥ 0 in order to achieve a better balance between strength and hot workability .72 is better, that is, it is desirable that the aluminum content is 2.6% by mass or less. More preferably, (0.6W _Ti + 0.31W _Nb + 0.15W _Ta )/ _WAl ≥ 0.93 should be satisfied for a better combination of strength and hot workability, and aluminum in the alloy must be limited to 2.3% by mass.

For the alloys of the present invention, high levels of tensile and creep properties are required at temperatures above 700°C. However, in combination with these properties, it is desirable to increase the oxidation resistance of the alloy. Oxidation can be a source of failure in some applications of the invention. Especially in high temperature, high stress applications in gas turbines, oxidation assisted cracking mechanisms can occur, for example, dwell fatigue can occur in turbine disk alloys. This cracking mechanism is promoted by rapid oxidation (and formation of unprotected oxides such as niobium oxide and titanium oxide) within the alloy when operating in extreme environments. Formation of a protective oxide helps to inhibit crack growth (by improving the oxidation resistance of the alloy). To reduce the susceptibility to oxidation cracking, the present invention increases the levels of elements that improve oxidation resistance (mainly chromium) and elements that can adversely affect oxidation (mainly chromium) that can form unprotected oxides. The purpose is to limit the use of titanium and niobium).

High levels of chromium are preferred to provide resistance to oxidation. Higher levels of chromium also improve resistance to hot corrosion. The chromium level should be at least 17.0% by weight. This provides better oxidation and corrosion resistance than other high strength alloys (TMW-4, Alloy 720Li, AD730, Rene 65). However, the addition of chromium increases the tendency of the alloy to form harmful TCP phases, such as the sigma (σ) phase. Limiting or stopping the precipitation of TCP phase formation is beneficial because TCP phases cause deterioration of material properties during high temperature operation. The property trade-off between oxidation resistance and high temperature strength is discussed below with reference to FIGS. 10-11.

Titanium is known to reduce the oxidation performance of chromia-forming superalloys by forming titanium oxide, which limits the protective capacity of the chromium-based oxide scale and accelerates the oxidation rate (S. Cruchley, et al., Chromia layer growth on a Ni-based superalloy: Sub-parabolic kinetics and the role of titanium, Corrosion Science 75:58-66, 2013). In the present invention, titanium is preferably limited to 3.5% by weight in order to provide better oxidation resistance. It is preferred to further reduce the level of titanium in order to provide a better level of oxidation resistance. Titanium is preferably limited to 3.0% by weight. This makes it possible to achieve oxidizability comparable to Waspaloy, which is known to have very good oxidizing properties. Waspaloy is also known to have excellent resistance to oxidation-assisted cracking mechanisms such as retention fatigue.

Similarly, the niobium content of the alloy should be limited to 3.0 wt% or less to protect against dwell fatigue. Niobium accelerates the growth of oxidation-assisted cracks when forming niobium oxide with a high pilling-to-Bedworth ratio (i.e., the large volume expansion of niobium oxide creates stress in the alloy as a result of the volume increase). It has been known. In embodiments particularly suitable for use in oxidizing environments, the niobium is 2.0% or less by weight. This is because by making the niobium content of the alloy 2.0% by mass or less, it shows little tendency to form niobium oxide when exposed to an oxidizing environment, that is, it does not adversely affect the alloy (Nemeth A.A.N. et. al., On the Influence of Nb/Ti Ratio on Environmentally Assisted Crack Growth in High-Strength Nickel Based Superalloys, Metal. And Mat Trans A 49:3923-3937, 2018).

The main alloying additions used to increase alloy strength are the gamma prime (γ') forming elements, aluminum, titanium, niobium and tantalum (determined in terms of strength merit index). Figures 5-8 show the effect of the elements aluminum, titanium and niobium on the strength of the alloy.

Based on FIGS. 5-8, the relationship between the amount of element added (based on titanium, aluminum and niobium) and the target alloy strength was derived. The relationships representing the three-dimensional aspects were determined as follows. The relationship between alloy strength and aluminum and titanium was determined from FIGS. 5-8 with a target strength of 1450 MPa or higher, which was determined at the isopleths of these values. The line translation as a function of different niobium contents was determined from FIGS. Based on the process described, the addition amounts of aluminum, titanium and niobium are found to follow the following equations (however, since niobium has a large effect on the APB energy intensity (see Equation 6), the effect on γ′ formation Note that 0.44 is applied as a factor for niobium instead of the 0.31 determined based on Equation 6, the APB strength factor for niobium is 21.4 and the factor applied for titanium is 15. Niobium is 1.42 times stronger in strength, so multiplying this factor by the aluminum equivalent of niobium (0.31) yields a factor of 0.44 (i.e., 1.42 x 0.31 = 0.44).
f(strength) = 0.6W _Ti + 0.44W _Nb + 0.2W _Al
where f(strength) is a number and W _Ti , W _Nb and W _Al are the weight percents of titanium, niobium and aluminum respectively in the alloy. In order to produce an alloy with a strength figure of merit of 1450 MPa or more, the numerical value of f(strength) must be 2.8 or more.

When the γ' volume fraction is 44% or less (f(γ') ≤ 4.6) and the maximum niobium content is 3.0 mass%, in order to achieve a strength merit index of 1450 MPa or more, titanium A minimum value of 1.8 mass % is required. The maximum niobium content of the alloy is therefore 3.0%. Preferably, the γ' volume fraction is limited to 40% or less (f(γ') ≤ 4.3), more preferably γ' volume fraction is 36% or less (f(γ') ≤ 4.0), i.e. , the titanium content is preferably 1.9% by weight or more and 2.0% by weight or more, respectively (FIG. 8). In order to achieve a strength figure of merit of 1450 MPa or more, the γ' volume fraction is 44% or less (f(γ') ≤ 4.6). A value of 2.3% by weight is required (FIG. 7). Preferably, the γ' volume fraction is limited to 40% or less (f(γ') ≤ 4.3), more preferably γ' volume fraction is 36% or less (f(γ') ≤ 4.0), i.e. , the titanium content is preferably 2.4% by weight or more and 2.5% by weight or more, respectively (FIG. 7).

With a maximum titanium content of 3.5% by weight, it is difficult to achieve a strength figure of merit of 1450 MPa without niobium in the alloy (FIG. 5). However, such alloys have exceptional hot workability due to their low γ' solvus. In the absence of niobium in the alloy, a minimum titanium content of 4.0 mass % achieves a strength figure of merit of 1450 MPa (FIG. 5). The maximum amount of titanium is therefore 4.0%. When titanium is 3.5 mass % or less, the addition of niobium can achieve a strength figure of merit of 1450 MPa. To achieve this strength figure of merit, an addition of niobium corresponding to 0.5% by weight of titanium is required. Therefore, to achieve the desired strength figure of merit of 1450 MPa, the alloy requires a minimum of 0.4 wt% niobium (0.5 wt% titanium divided by a factor of 0.6 and multiplied by a factor of 0.44, determined by corresponds to 0.4% by weight of niobium). More preferably, titanium is limited to 3.0 wt%, which means that 1.0 wt% equivalent niobium is required to achieve a target strength figure of merit of 1450 MPa. is 0.7% by weight niobium. Most preferably, if titanium is limited to 3.0 wt.% and a strength figure of merit of 1550 MPa is to be achieved, 1.6 wt.% titanium equivalent niobium is required, which means a niobium content of 1.2 wt.% or more. corresponds to

Tantalum is any element present in an amount less than or equal to 3.05% by weight. This limit is to control alloy costs (see Figure 12). In one embodiment of the present invention, tantalum is added to replace niobium to a level of 3.05% by weight, and based on tantalum's density (.15/.31) and strength factor (27.1/21.4), 0.5%. A replacement factor of 66 Ta is derived. Even with 3.05 mass % tantalum and no niobium in the alloy, an alloy with a high strength figure of merit can be achieved. Preferably, tantalum is limited to 2.25 wt%, more preferably to 1.5 wt%, most preferably to 0.65 wt% to control costs (see Figure 12). described later).
f(strength) = 0.6 W _Ti + 0.44 (W _Nb + 0.66 W _Ta ) + 0.2 W _Al
In order to produce an alloy with a strength figure of merit of 1450 MPa or more, the value of f(strength) must be 2.8 or more.

It is desirable to design alloys with a strong gamma matrix phase, along with an increase in strength contribution due to precipitation hardening due to the gamma prime phase (calculated in terms of the strength figure of merit). The strength of the gamma matrix can be calculated in terms of the solid solution figure of merit. Solid solution strength is particularly important for imparting high temperature strength (both tensile and creep strength) to the alloy. Tungsten and molybdenum have large coefficients and strongly partition to the gamma phase. In contrast, niobium and tantalum have large intensity coefficients but limited partitioning into the gamma phase (see Eq. 10). Therefore, the solid solution strength of the gamma phase of the alloy mainly depends on the amount of tungsten and molybdenum added. FIG. 9 shows the relationship between molybdenum, tungsten and solid solution index. Since the density of tungsten is about twice that of molybdenum, a factor of 0.5 is applied to the tungsten content. That is, this factor accounts for the difference in density. A solid solution index of 97 MPa or greater (comparable to TMW-4) is desirable to achieve creep strength equal to or greater than the alloys listed in Table 3 while maintaining excellent levels of workability. . From FIG. 10, the relationship between the amount of tungsten and molybdenum added and the solid solution index can be determined.
f(solid solution)= _WMo + _0.5WW

Here, in order to realize a desirable alloy having a solid solution figure of merit of 95 MPa or more, f (solid solution) must be a numerical value of 3.4 or more, and W _Mo and W _W are molybdenum in the alloy, respectively. and mass % of tungsten. It is desirable that f(solid solution) is 3.6 or more. Thereby, the solid-solution merit index can be made 100 MPa or more.

FIG. 10 shows the γ′ volume fraction and the sum of the elements molybdenum and tungsten (based on the relationship W _Mo +0.5 _{W W} ) at the maximum achievable chromium content at a target stability number of 0.9; shows the influence of A correction factor of 0.5 is applied to tungsten because it has about twice the density of molybdenum. This factor indicates a substantial difference in the density of the elements.

Based on the strength requirements of the alloy (in terms of strength figure of merit), the γ' volume fraction is 29% or more. In order to achieve creep resistance and high-temperature strength, the solid solution index must be high, and the minimum value of (W _Mo +0.5 _{W W} ) must be 3.4 mass %. Based on these minimum requirements and the need for a chromium content of at least 17.0% by weight, the γ′ volume fraction should be limited to 44%. At the level of γ' volume fraction (29%), W _Mo +0.5W _W is limited to 6.3 wt%. Therefore, the maximum Mo content is 6.3% or less. The maximum chromium content of the alloy is limited to 21.3% by mass or less to ensure a balance between strength (γ' volume fraction and high temperature strength), (W _Mo +0.5 W _W )) and oxidation resistance ( Figure 10). Preferably, the chromium content of the alloy is greater than or equal to 18.0% by weight, which provides the alloy with better oxidation and corrosion resistance. At this level of chromium, the volume fraction of the γ' phase is preferably limited to 40% and the sum of elemental molybdenum and elemental tungsten is limited to 5.6% by weight (according to the relationship Mo+0.5W). be.

Preferably, a target stability of 0.89 or less is desirable to ensure a better level of microstructural stability and avoid TCP formation. The alloys with the most desirable stability in Table 3 have a stability index number of 0.89 or less. FIG. 11 shows the γ′ volume fraction and the sum of the elements molybdenum and tungsten (based on the relationship W _Mo +0.5 _{W W} ) for the maximum achievable chromium content at the target stability number of 0.89. Show impact. For a stability number of 0.89, the chromium content in the alloy is preferably limited to 19.5% by weight or less. The maximum volume fraction of the γ' phase is preferably limited to 39% or less. Preferably, the sum of the elements molybdenum and tungsten (based on the relationship W _Mo +0.5W _W ) is limited to 5.1% by weight when the preferred chromium content (18% by weight or more) is included in the alloy. be. At this level of chromium, the volume fraction of the γ' phase is limited to 36% and the sum of elemental molybdenum and tungsten is limited to 4.3 mass (based on the relationship _WMo + _0.5WW ). .

Among the cast and forged alloys defined in Table 1, alloy AD730 is the lowest cost alloy suitable for high stress applications at operating temperatures above 700°C. The cost of the alloy based on current element prices (2019) is estimated at $14.6/kg. The most popular cast and wrought alloy for these applications is Alloy 718, and an example of the composition of Alloy 718 is given (Ni-19Cr-0.8Ti-0.5Al-3Mo-29.7Fe-5Nb-0.06C), The estimated cost of this alloy is $9.9/kg. However, the IN718 is limited to operating temperatures below 650°C in high stress applications. A significantly lower target cost ($14/kg) than the AD730 is desired for the present invention. In the compositional region defined in Table 2, the main elements that affect cost are tantalum and cobalt, and the effect of these elements on alloy cost is shown in FIG. Therefore, it is desirable to have a maximum cobalt content of 7.1 wt%, and preferably to limit cobalt to 5.3 wt% or less, as alloy costs are limited to $13/kg. More preferably, the alloy cost is limited to $12/kg, so it is preferable to limit cobalt to 3.5 wt% or less, and most preferably, the alloy cost is limited to $11/kg. , and cobalt is most preferably limited to 1.6% by weight or less.

From FIG. 12, it can be seen that the added amounts of cobalt and tantalum follow the following formula.
f(cost) = W _Ta + 0.43 W _Co
where f(cost) is a number and W _Co is the weight percent of cobalt in the alloy. In order to produce an alloy with a cost of 14$/kg or less, the value of f(cost) must be 3.05 or less. To produce alloys with elemental costs of 13$/kg or less, 12$/kg or less, and 11$/kg or less, f(cost) values should be 2.25 or less, more preferably 1.5 or less, and most preferably 0.65, respectively. It is desirable to:

Iron can be added to the alloy to reduce costs and increase the recyclability of the alloy. Adding iron can destabilize the microstructure. Limiting the iron addition to 10.0 wt% produces a good balance of low cost, improved recyclability and microstructural stability. More preferably, the content of iron is limited to 6% by mass or less, and more preferably within the range of 1.0% by mass to 5.0% by mass. This provides the best balance between cost, recyclability and alloy performance.

Tungsten and tantalum elements can greatly improve the strength and creep resistance of the alloy. However, since these elements are much denser than nickel, the alloy can also be made denser. It is important to add these elements in a manner that balances increased strength with increased density. FIG. 13 shows the effect of these elements on the alloy density for alloys with a γ' phase volume fraction of 29-44%. A target density of 8.6 g/cm ³ or less is desirable to avoid higher densities than the alloys listed in Table 3. Therefore, the amount of tungsten added is limited to 4.9% by mass or less. As mentioned earlier, tungsten also plays a role in providing strength and creep resistance through solid-solution strengthening, but to achieve the required solid-solution strengthening, it was previously determined that Mo+0.5W ≥ 3.4 wt%. there is Molybdenum is added at least 0.9% by weight based on the maximum tungsten addition (4.9% by weight). More preferably, tungsten is limited to 2.7 wt% or less to achieve a density of 8.5 g/ ^cm3 or less, and most preferably tungsten is limited to 8.4 g/ ^cm3 or less to achieve a density of is limited to 0.7% by mass or less. Therefore, it is preferred to have a minimum molybdenum content of at least 2.0 wt%, most preferably 3.0 wt%.

Additions of carbon, boron and zirconium are necessary to give strength to the grain boundaries. This is particularly beneficial for the creep and fatigue properties of the alloy. The carbon concentration should be limited to 0.1 wt% or less, more preferably 0.07 wt% or less, as carbon improves the forgeability of the alloy. Boron concentration is limited to 0.001-0.1% by weight, preferably 0.03% by weight or less, since boron can separate into the liquid phase during solidification and cause liquid cracking during welding of the alloy. More preferably, the boron concentration is limited to 0.02 mass % or less to ensure good weldability. The zirconium concentration should be in the range of 0.001 wt% to 0.5 wt%, preferably 0.01 wt% or less, more preferably 0.006 wt% or less. The addition of magnesium can be used to improve the hot ductility of wrought alloys, improving hot workability and creep rupture life. Magnesium additions up to 200 PPM are desirable, preferably between 10 PPM and 100 PPM magnesium additions to the alloy to improve hot workability and creep rupture properties.

When the alloy is manufactured, it is beneficial that the alloy be substantially free of unavoidable impurities. These impurities may include elemental sulfur (S), elemental manganese (Mn) and elemental copper (Cu). Elemental sulfur should be 0.003% by mass (30 PPM in terms of mass) or less. Manganese is an unavoidable impurity limited to 0.25% by weight, preferably 0.1% by weight or less. Copper (Cu) is an unavoidable impurity preferably limited to 0.5% by weight. The presence of sulfur above 0.003 wt% can cause embrittlement of the alloy, and sulfur segregates at the alloy/oxide interface formed during oxidation. For this reason, the sulfur level is preferably 0.001 wt% or less. Vanadium, an unavoidable impurity, is preferably limited to 0.5% by weight, preferably 0.3% by weight or less, most preferably 0.1% by weight or less, because it adversely affects the oxidation behavior of the alloy. be done. This segregation can lead to increased spallation of the protective oxide scale. If the concentration of these unavoidable impurities exceeds the specified levels, problems with product yield and deterioration of the material properties of the alloy are to be expected.

Hafnium (Hf) is added up to 0.5% by mass, more preferably up to 0.2% by mass, to bind unavoidable impurities in the alloy and to provide strength. That is useful. Hafnium is a strong carbide former and can lead to further grain boundary strengthening.

So-called "reactive elements" (yttrium (Y), lanthanum (La) and cerium (Ce)) are added at levels up to 0.1% by weight. This is beneficial for improving the adhesion of protective oxide _layers such as _Cr2O3 . These reactive elements can "sweep" harmful elements such as sulfur. Sulfur segregates at the alloy oxide interface and weakens the bond between the oxide and the substrate, resulting in exfoliation of the oxide. Silicon (Si) can be beneficially added up to 0.5% by weight. Additions of silicon to nickel-based superalloys at levels up to 0.5% by weight have been shown to be beneficial to oxidation properties. Silicon in particular segregates at the alloy/oxide interface and improves the bonding strength of the oxide to the substrate. This suppresses exfoliation of the oxide, resulting in improved oxidation resistance.

Based on the description of the invention in this section, Table 4 lists the broad scope of the invention. Table 4 also shows preferred ranges. Table 4 shows the composition range in mass % for the new design alloys.

The following section describes exemplary compositions of the invention. Calculated properties for these new alloys are listed. The rationale for the design of these alloys is now explained.

Three examples of the invention are listed in Table 5. The predicted properties of this alloy are listed in Table 6. The alloys listed were determined by lowering the γ′ solvus temperature compared to the prior art alloys listed in Table 3 (especially other high strength alloys with strength merit figures of 1450 MPa or more). The advantage is that an improved combination of hot workability and high strength (in terms of strength figure of merit) is achieved. These improvements in material performance are achieved while maintaining lower elemental costs than alloys with strength indices of 1450 MPa or greater. Table 5 is an example of an alloy of the present invention compared to the alloy examples listed in Table 1.

A comparison of the alloys of the invention with the prior art shows that the alloys of the invention have much lower levels of cobalt. Cobalt is commonly added to prior art alloys as an element that lowers the gamma prime solvus. Most of the prior art determines the preferred range of cobalt based on its technical effect (increasing cobalt reduces solvus and thus improves processability). On the other hand, the alloys described herein achieve low gamma prime solvus at low cobalt content by balancing other elements, such as gamma prime forming elements. This results in an alloy with good hot workability and low elemental cost. Table 6 shows the calculated phase fractions and figures of merit produced by the "Alloy Design" software. Table 6 shows the results for the compositions listed in Table 5 and the example alloys listed in Table 1.

Claims (27)

17.0-21.3% Chromium, 7.1% Cobalt, 0.9-6.3% Molybdenum, 4.9% Tungsten, 1.8-3.2%, by weight aluminum, 1.8-4.0% titanium, ≤3.05% tantalum, ≤3.0% niobium, ≤0.1% carbon, 0.001-0.1% boron, 0 .001-0.5% zirconium, 0.02% or less magnesium, 0.5% or less silicon, 0.1% or less yttrium, 0.1% or less lanthanum, 0.1% or less cerium, 0.003% or less sulfur, 0.25% or less manganese, 0.5% or less copper, 0.5% or less hafnium, 0.5% or less vanadium, and 10.0% or less iron, A nickel-based alloy composition, the balance being nickel and incidental impurities. 2. The nickel-based alloy composition according to claim 1, satisfying the following formulae, where W _Al , W _Ti , W _Nb and W _Ta are the weight percentages of aluminum, titanium, niobium and tantalum contained in the alloy.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≤ 1.1
Preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≤ 1.0 3. The nickel-based alloy composition according to claim 1 or 2, wherein the mass percentages of aluminum, titanium, niobium and tantalum contained in the alloy are W _Al , W _Ti , W _Nb and W _Ta , respectively.
(0.6W _Ti + 0.31W _Nb + 0.15W _Ta )/ _WAl ≥ 0.64
Preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.72
More preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.76
Even more preferably, the following formula is satisfied.
(0.6 W _Ti + 0.31 W _Nb + 0.15 W _Ta )/W _Al ≥ 0.93 The volume fraction of the gamma prime phase contained in the alloy is 48% or less at 760°C, preferably 44% or less at 760°C, more preferably 40% or less at 760°C, most preferably 760°C. 4. The nickel-based alloy composition according to any one of claims 1 to 3, wherein the total content of the nickel-base alloy composition is 36% or less. 5. The nickel according to any one of claims 1 to 4, which satisfies the following formulas, where W _Al , W _Ti , W _Ta and W _Nb are mass percentages of aluminum, titanium, tantalum and niobium contained in the alloy. Base alloy composition.
3.4 ≤ 0.6W _Ti + 0.31W _Nb + 0.15W _Ta + 0.79W _Al
Preferably, the following formula is satisfied.
3.6 ≤ 0.6W _Ti + 0.31W _Nb + 0.15W _Ta + 0.79W _Al 6. The nickel according to any one of claims 1 to 5, which satisfies the following formulas, where the mass % of aluminum, titanium, tantalum and niobium contained in the alloy is W _Al , W _Ti , W _Ta and W _Nb respectively. Base alloy composition.
4.6≧0.6W _Ti +0.31W _Nb +0.15W _Ta +0.79W _Al
Preferably, the following formula is satisfied.
4.3≧0.6W _Ti +0.31W _Nb +0.15W _Ta +0.79W _Al
More preferably, the following formula is satisfied.
4.0≧0.6W _Ti +0.31W _Nb +0.15W _Ta +0.79W _Al 7. The nickel-base alloy composition according to any one of claims 1 to 6, comprising 18.0% or more of chromium by mass. 8. A nickel-base alloy composition according to any one of claims 1 to 7, comprising 19.5% or less of chromium by mass. 9. A nickel-base alloy composition according to any one of the preceding claims, comprising 2.25% or less, preferably 1.5% or less, most preferably 0.65% or less of tantalum, by mass. 10. A nickel-base alloy composition according to any one of the preceding claims, comprising molybdenum in mass % of 2.0% or more, preferably 3.0% or more. 1.9% or more, preferably 2.0% or more, more preferably 2.3% or more, even more preferably 2.4% or more, and most preferably 2.5% or more, by mass % of titanium; A nickel-based alloy composition according to any one of claims 1-10. 12. The nickel-base alloy composition according to any one of claims 1 to 11, comprising no more than 0.2% by weight of hafnium. 13. A nickel-base alloy composition according to any one of the preceding claims, comprising tungsten in mass % of 2.7% or less, preferably 0.7% or less. 14. The nickel-base alloy composition according to any one of claims 1 to 13, comprising 2.0% or less of niobium by mass. 15. A nickel-base alloy composition according to any one of the preceding claims, comprising titanium in mass % of 3.5% or less, preferably 3.0% or less. 16. The nickel-base alloy composition according to any one of claims 1 to 15, comprising 1.9% or more, by mass, of aluminum. 17. A nickel-base alloy composition according to any one of the preceding claims, comprising aluminum in weight percent of 2.75% or less, preferably 2.6% or less, more preferably 2.3% or less. 18. A nickel-based alloy composition according to any one of the preceding claims, comprising cobalt in mass % of 5.3% or less, preferably 3.5% or less, more preferably 1.6% or less. 19. A nickel-base alloy composition according to any one of the preceding claims, comprising vanadium in % by weight of 0.3% or less, preferably 0.1% or less. 20. The nickel-based alloy composition according to any one of claims 1 to 19, which satisfies the following formulas, where W _Ta and W _Co are the mass percentages of tantalum and cobalt contained in the alloy, respectively.
W _Ta +0.43 W _Co ≦3.05
Preferably, the following formula is satisfied.
W _Ta +0.43 W _Co ≦2.25
More preferably, the following formula is satisfied.
W _Ta +0.43 W _Co ≦1.5
Most preferably, the following formula is satisfied.
W _Ta +0.43 W _Co ≦0.65 21. Nickel according to any one of claims 1 to 20, which satisfies the following formulas, where W _Ti , W _Nb , W _Ta and W _Al are the weight percentages of titanium, niobium, tantalum and aluminum contained in the alloy, respectively: Base alloy composition.
0.6W _Ti + 0.44 ( _WNb + 0.66W _Ta ) + 0.2W _Al ≥ 2.8 0.3% or more, preferably 0.4% or more, more preferably 0.7% or more, still more preferably 0.8% or more, and most preferably 1.2% or more of niobium in mass%; 22. The nickel-based alloy composition of any one of claims 1-21. 23. The nickel-based alloy composition according to any one of claims 1 to 22, wherein the composition satisfies the following formulas, where W _Mo and W _W are the mass percentages of molybdenum and tungsten contained in the alloy, respectively.
W _Mo +0.5 _{W W} ≧3.4
Preferably, the following formula is satisfied.
W _Mo +0.5 _{W W} ≧3.6 24. The nickel-based alloy composition according to any one of claims 1 to 23, wherein the composition satisfies the following formulas, where W _Mo and W _W are the mass percentages of molybdenum and tungsten contained in the alloy, respectively.
_WMo + _0.5WW ≦6.3
Preferably, the following formula is satisfied.
_WMo + _0.5WW≤5.6
More preferably, the following formula is satisfied.
_WMo + _0.5WW≤5.1
Most preferably, the following formula is satisfied.
_WMo + _0.5WW≤4.3 Claims 1 to 24, wherein the volume fraction of gamma prime phase contained in the alloy is at least 29% at 760°C, preferably at least 31% at 760°C and most preferably at least 40% at 760°C. The nickel-based alloy composition according to any one of 26. A nickel-base alloy composition according to any one of the preceding claims, comprising no more than 6.0%, preferably no more than 5.0% iron, by mass. Cast and forged parts formed from the nickel-based alloy composition of any one of claims 1-26.

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