The present invention relates to a nickel-based superalloy composition for use as a turbine disc component within a gas turbine engine and other turbomachinery. The turbine disc is a critical component in gas turbine engines. Increases in turbine disc alloy performance - in terms of maximum operating temperature and maximum service - life can have a significant impact on the efficiency of the engine as well as the cost effectiveness of operating the engine.

Examples of typical compositions of nickel-based superalloys which are used for turbine disc components are listed in Table 1. In development of higher strength alloys there has been a tendency to move towards higher levels of cost, Figure 1. Table 1: Nominal composition in wt.% of commonly applied nickel-based superalloys used for powder metallurgy turbine discs.

Alloy (wt.%) Cr Co Fe Mo W Al Ti Ta Nb Hf C B Zr

N18 11.50 15.70 0.00 6.50 0.00 4.35 4.35 0.00 0.00 0.50 0.015 0.015 0.030

Rene88DT 16.00 13.00 0.00 4.00 4.00 2.10 3.70 0.00 0.70 0.00 0.030 0.015 0.030

RR1000 15.00 18.50 0.00 5.00 0.00 3.00 3.60 2.00 0.00 0.50 0.030 0.020 0.060

ME3 13.00 20.50 0.00 3.70 2.00 3.40 3.60 2.40 0.90 0.00 0.040 0.030 0.050

AlloylO 10.46 17.93 0.00 2.52 4.74 3.53 3.79 1.61 0.97 0.00 0.027 0.028 0.070

LSHR 13.00 20.50 0.00 2.75 4.55 3.50 3.50 1.70 1.55 0.00 0.030 0.030 0.050

US8,147,749 10.09 19.60 0.00 2.79 2.62 3.14 2.17 7.28 1.55 0.40 0.030 0.030 0.050

N19 12.90 11.80 0.00 4.70 3.20 2.50 3.80 0.00 1.60 0.30 0.022 0.015 0.060

US8,613,810 12.00 18.00 0.00 2.90 2.80 3.20 3.10 5.10 1.50 0.40 0.055 0.025 0.055

US2015/0192002 13.40 14.00 0.00 1.30 5.30 3.25 3.10 3.70 1.55 0.00 0.050 0.025 0.055

It is the aim of the present invention is deliver substantially high strength in combination with a reduction in alloy cost and an improvement in oxidation/corrosion resistance. The balance of properties for the new alloy make it more cost effective for the production of components for high temperature applications; in particular for use as a turbine disc application where the operating temperature for the component is 800°C or greater.

The present invention provides a nickel-based alloy composition consisting, in weight percent, of: a nickel -based alloy composition consisting, in weight percent, of: between 9.5 and 14.4% chromium, between 3.8 and 9.6% cobalt, between 0.0 and 4.0%) iron, between 0.6 and 5.1% molybdenum, between 0.7 and 6.2% tungsten, between 2.6 and 3.3%> aluminium, between 2.3 and 5.6% titanium , between 0.0 and 4.0% niobium, between 0.0 and 2.4% tantalum, between 0.01 and 0.1%) carbon, between 0.001 and 0.1% boron, between 0.001 and 0.3% zirconium, between 0.0 and 0.5%) silicon, between 0.0 and 0.1% yttrium, between 0.0 and 0.1% lanthanum, between 0.0 and 0.1% cerium, between 0.0 and 0.003%) sulphur, between 0.0 and 0.25% manganese, between 0.0 and 0.5%> vanadium, between 0.0 and 0.5% copper, and between 0.0 and 0.5% hafnium, the balance being nickel and incidental impurities, wherein the following equation is satisfied in which WNb, WT ₃ , WTI and WAI are the weight percent of niobium, tantalum, titanium and aluminium in the alloy respectively

0.6W _Tl + 0.31Η^ _ϋ + 0.27W _Ta + 0.12W _A1 > 3.7.

This alloy provides a good compromise between high strength, cost, oxidation/corrosion resistance and resistance to TCP formation.

In a preferred embodiment, the following equation is satisfied in which WNb, WT ₃ and Wco are the weight percent of niobium, tantalum and cobalt in the alloy respectively,

W _m + l.S3W _Co + 3.7W _Ta ≤ 14.7

preferably W _m + 1.53W _Co + 3.7W _Ta ≤ 12.4

most preferably W _nb + 1.53W _CO + 3.7W _Ta ≤ 9.5.

Such alloys are optimised for reduced cost. In a preferred embodiment the following equation is satisfied in which WNb, Wia, W-n and WAI are the weight percent of niobium, tantalum, titanium and aluminium in the alloy respectively

0.6W _Ti + 0.311½, + 0.27W _Ta + 0.12W _A1 > 4.0.

Such alloys have even greater strength.

In an embodiment the following equation is satisfied in which WNb, Wia, W-n and WAI are the weight percent of niobium, tantalum, titanium and aluminium in the alloy respectively

^' (0.6 ^Ti+0.31 0¼+0.15 Wi _a )IW _M <1.3.

Such alloys have high γ' phase stability and resist unwanted eta phase and delta phase. Thus such alloys have good ductility and fatigue resistance.

In an embodiment the following equation is satisfied in which WNb, WT ₃ , W-n and WAI are the weight percent of niobium, tantalum, titanium and aluminium in the alloy respectively

0.6WT, + Q.31W _Nh + 0.15W _Ta + 0.914// preferably 0.6W _Ti + Q.31W _Nb + 0.1SW _Ta + 0.9W _M ≤ 5.9 such alloys have a solvus temperature of the γ' phase of less than 1180°C allowing heat treatment above the gamma-prime solvus whilst reducing susceptibility of the alloy to cracking on cool down from above the γ' solvus temperature. In an embodiment the following equation is satisfied in which WNb, Wia, WTI and WAI are the weight percent of niobium, tantalum, titanium and aluminium in the alloy respectively

5.4 < 0.6I ½ + 0.31W _Nb + 0.15W _Ta + 0.8W _M ≤ 6.0

Such an alloy has good strength properties.

In an embodiment the following equation is satisfied in which Ww and Wia, are the weight percent of tungsten and tantalum, in the alloy respectively

W _w + l.lW _Ta ≤ 6.2 preferably W _w + l.lW _Ta ≤ 4.6 most preferably W _w + l.lW _Ta ≤ 2.9

Such an alloy has limited density whilst maintaining a high strength. In an embodiment the following equation is satisfied in which Ww, Wco and WMO, are the weight percent of tungsten, cobalt and molybdenum, in the alloy respectively

W _w + .96W _CO + 1.2W _Mo ≥ 16.0

Such an alloy has particularly high creep resistance.

In an embodiment the nickel-based alloy composition consists of in weight percent, 10.0 wt.% or more chromium, more preferably 1 1.5 wt.% or more chromium. Such an alloys has improved corrosion resistance.

In an embodiment the nickel -based alloy composition consists of in weight percent, 5.5 wt.% or more cobalt. Such an alloy has an improved balance of creep resistance and alloy stability.

In an embodiment the nickel -based alloy composition consists of, weight percent, 8.1 wt.% or less cobalt, preferably 8.0 wt.% or less cobalt, more preferably 6.0 wt.% or less cobalt. Such an alloy has reduced cost. In an embodiment the nickel-based alloy composition consists of in weight percent, 1.0 wt.% or more iron. Such an alloy has the benefit of being less expensive and is more easily recycled.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 2.0 wt.%) or less iron. Such an alloy has improved microstructural stability whilst providing a good balance of low cost and improved recyclability.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 2.2 wt.% or more molybdenum, preferably 3.6 wt.% or more molybdenum. Such an alloy has a better balance of creep resistance and density. In an embodiment the nickel-based alloy composition consists of, in weight percent, 4.1 wt.% or less molybdenum. Such an alloy has still further improved alloy stability.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 1.9 wt.%) or more tungsten. Such an alloy has an even better balance of creep resistance and alloy stability.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 4.6 wt.%) or less tungsten, preferably 2.9 wt.%> or less tungsten. This results in an alloy with an even lower density.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 2.6 wt.%) or more aluminium. Such an alloy achieves an improved combination of strength, γ' solvus and stability of the γ'.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 3.0 wt.%) or less aluminium. Such an alloy achieves a γ' solvus at a lower temperature.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 2.6 wt.%) or more titanium, preferably 3.0 wt.%> or more titanium, more preferably 3.5 wt.%> or more titanium, most preferably 4.1 wt.%> or more titanium. Increasing the titanium content enables a reduction in the amount of tantalum needed to provide an alloy of a certain strength and this reduces cost for such an alloy.

In an embodiment the nickel -based alloy composition consists of, in weight percent, 1.8 wt.% or less tantalum, preferably 1.0 wt.% or less tantalum. Such an alloy has reduced cost and density.

In an embodiment the nickel-based alloy composition consists of, in weight percent, 2.7 wt.%) or less niobium. This achieves an alloy with improved creep properties at reduced cost. In an embodiment the nickel-based alloy composition consists of 51 - 58% volume fraction gamma-prime at 850°C. This achieves an alloy with a high degree of strength in combination with a lower gamma-prime solvus.

In an embodiment of the present invention a turbine disc is provided which is formed of the nickel-based alloy composition of the present invention.

In an embodiment of the invention a gas turbine engine comprises the turbine disc made from the nickel-based alloy composition of the present invention.

The term "consisting of is used herein to indicate that 100% of the composition is being referred to and the presence of additional components is excluded so that percentages add up to 100%. Unless otherwise stated, percent's are expressed in weight percent.

The invention will be more fully described, by way of example only, with reference to the accompanying drawings in which:

Figure 1 shows the calculated trade-off between the elemental cost and yield strength (in terms of strength merit index) for the alloys listed in Table 1 and alloys within the alloy design space listed in Table 2;

Figure 2 is a contour plot showing the effect of cobalt and niobium on the elemental cost of alloys when the tantalum content is fixed at 0.0 wt.%;

Figure 3 is a contour plot showing the effect of cobalt and niobium on the elemental cost of alloys when the tantalum content is fixed at 1.0 wt.%; Figure 4 is a contour plot showing the effect of cobalt and niobium on the elemental cost of alloys when the tantalum content is fixed at 2.0 wt.%;

Figure 5 is a contour plot showing the effect of cobalt and niobium on the elemental cost of alloys when the tantalum content is fixed at 3.0 wt.%;

Figure 6 is a contour plot showing the effect of cobalt and niobium on the elemental cost of alloys when the tantalum content is fixed at 4.0 wt.%;

Figure 7 is a contour plot showing the effect of elements aluminium and titanium on yield strength (in terms of strength merit index), for alloys with the preferred elemental cost (<14.7 $/kg) when the niobium content is fixed at 0.0 wt.% and tantalum content is fixed at 0.0 wt.%. Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship (0.6#¾+0.31 W¾b)/^Ai; Figure 8 is a contour plot showing the effect of elements aluminium and titanium on yield strength (in terms of strength merit index), for alloys with the preferred elemental cost (<14.7 $/kg) when the niobium content is fixed at 1.0 wt.% and tantalum content is fixed at 0.0 wt.%. Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship (0.6 Wn+0.31 Wm)/WAi;

Figure 9 is a contour plot showing the effect of elements aluminium and titanium on yield strength (in terms of strength merit index), for alloys with the preferred elemental cost (<14.7 $/kg) when the niobium content is fixed at 2.0 wt.% and tantalum content is fixed at 0.0 wt.%). Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship W b)/ WA\',

Figure 10 is a contour plot showing the effect of elements aluminium and titanium on yield strength (in terms of strength merit index), for alloys with the preferred elemental cost (<14.7 $/kg) when the niobium content is fixed at 3.0 wt.% and tantalum content is fixed at 0.0 wt.%). Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship (0.6 Wn+0.31 Wm)/WA\,'

Figure 11 is a contour plot showing the effect of elements aluminium and titanium on yield strength (in terms of strength merit index), for alloys with the preferred elemental cost (<14.7 $/kg) when the niobium content is fixed at 4.0 wt.% and tantalum content is fixed at 0.0 wt.%). Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship (0.6 Wn+0.31 Wm /WA ;

Figure 12 is a contour plot showing the effect of elements aluminium and niobium plus titanium (according to the relationship 0.6Wn+0.3 Wwb) on γ' solvus temperature, for alloys with the preferred elemental cost (<14.7 $/kg) when tantalum content is fixed at 0.0 wt.%>. Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship (0.6 Wn+0.31 and the preferred concentration of aluminium titanium and niobium according to the relationship (0.6Wn+0.3 Wm+0A2WAi=3 );

Figure 13 is a contour plot showing the effect of elements aluminium and niobium plus titanium (according to the relationship 0.6Wn+0.3 W b) on volume fraction of γ' at 900°C, for alloys with the preferred elemental cost (<14.7 $/kg) when tantalum content is fixed at 0.0 wt.%). Superimposed is the preferred limits for the ratio of the elements titanium, niobium and aluminium according to the relationship (0.6Wn+0.31 Wm)/WAi and the preferred concentration of aluminium titanium and niobium according to the relationship (0.6f _T i+0.3 «¾b+0.12^Ai=3.7);

Figure 14 is a contour plot showing the effect of elements tungsten and tantalum on alloy density; Figure 15 is a contour plot showing the effect of elements tungsten and molybdenum on creep resistance (in terms of creep merit index) when the tantalum content is fixed at 0.0 wt.%;

Figure 16 is a contour plot showing the effect of elements tungsten and molybdenum on creep resistance (in terms of creep merit index) when the tantalum content is fixed at 1.0 wt.%;

Figure 17 is a contour plot showing the effect of elements tungsten and molybdenum on creep resistance (in terms of creep merit index) when the tantalum content is fixed at 2.0 wt.%;

Figure 18 is a contour plot showing the effect of elements tungsten and molybdenum on creep resistance (in terms of creep merit index) when the tantalum content is fixed at 3.0 wt.%;

Figure 19 is a contour plot showing the effect of elements tungsten and molybdenum on creep resistance (in terms of creep merit index) when the tantalum content is fixed at 4.0 wt.%; Figure 20 is a contour plot showing the effect of elements tungsten and molybdenum on creep resistance (in terms of creep merit index) when the tantalum content is fixed at 5.0 wt.%;

Figure 21 is a contour plot showing the effect of elements chromium and tungsten plus molybdenum (according to the relationship WMO+0.5WW) on the stability number Md. Traditionally, nickel-based superalloys have been designed through empiricism. Thus their chemical compositions have been isolated using time consuming and expensive experimental development, involving small-scale processing of limited quantities of material and subsequent characterisation of their behaviour. The alloy composition adopted is then the one found to display the best, or most desirable, combination of properties. The large number of possible alloying elements indicates that these alloys are not entirely optimised and that improved alloys are likely to exist.

In superalloys, generally additions of chromium (Cr) and aluminium (Al) are added to impart resistance to oxidation/corrosion, cobalt (Co) is added to improve resistance to sulphidisation. For creep resistance, molybdenum (Mo), tungsten (W) and cobalt (Co) are introduced, because these retard the thermally-activated processes - such as, dislocation climb - which determine the rate of creep deformation. To promote static and cyclic strength, aluminium (Al), tantalum (Ta), niobium (Nb) and titanium (Ti) are introduced as these promote the formation of the precipitate hardening phase gamma-prime (γ'). This precipitate phase is coherent with the face-centered cubic (FCC) matrix phase which is referred to as gamma (y).

A modelling-based approach used for the isolation of new grades of nickel-based superalloys is described here, termed the "Alloys-By-Design" (ABD) method. This approach utilises a framework of computational materials models to estimate design relevant properties across a very broad compositional space. In principle, this alloy design tool allows the so called inverse problem to be solved; identifying optimum alloy compositions that best satisfy a specified set of design constraints.

The first step in the design process is the definition of an elemental list along with the associated upper and lower compositional limits. The compositional limits for each of the elemental additions considered in this invention - referred to as the "alloy design space" - are detailed in Table 2.

Table 2: Alloy design space studied.

Alloy (wt.%) Cr Co Mo W Al Ti Ta Nb c B Zr

Min 6.00 0.00 0.00 0.00 2.00 0.00 0.00 0.00

0.050 0.025 0.055

Max 18.00 24.00 8.00 8.00 5.00 8.00 10.00 4.00

The second step relies upon thermodynamic calculations used to calculate the phase diagram and thermodynamic properties for a specific alloy composition. Often this is referred to as the CALPHAD method (CALculate PHAse Diagram). These calculations are conducted at the typical service temperature for the new alloy (900°C), providing information about the phase equilibrium (microstructure).

A third stage involves isolating alloy compositions which have the desired microstructural architecture. In the case of nickel based superalloys which require superior resistance to creep deformation, the creep rupture life generally improves as the volume fraction of the precipitate hardening phase γ' is increased, the most beneficial range for volume fraction of γ' lies between 60%-70% at 900°C (however often due to other design restraints volume fraction may be limited to lower values than this and so alloys with a γ' volume fraction of 50% to 60% are included). At values above 70%) volume fraction of γ' a drop in creep resistance is observed.

It is also necessary that the γ/γ' lattice misfit should conform to a small value, either positive or negative, since coherency is otherwise lost; thus limits are placed on its magnitude. The lattice misfit δ is defined as the mismatch between γ and γ' phases, and is determined according to

2( . - a _v )

a _f + a _y where α _γ and ci are the lattice parameters of the γ and γ' phases.

Thus the model isolates all compositions in the design space which are calculated to result in a desired volume fraction of γ', which have a lattice misfit γ' of less than a predetermined magnitude.

In the fourth stage, merit indices are estimated for the remaining isolated alloy compositions in the dataset. These include: creep-merit index (which describes an alloy's creep resistance based solely on mean composition), strength-merit index (which describes an alloy's precipitation yield strength based solely on mean composition), density, cost, stable microstructure and gamma-prime solvus temperature.

In the fifth stage, the calculated merit indices are compared with limits for required behaviour, these design constraints are considered to be the boundary conditions to the problem. All compositions which do not fulfil the boundary conditions are excluded. At this stage, the trial dataset will be reduced in size quite markedly.

The final, sixth stage involves analysing the dataset of remaining compositions. This can be done in various ways. One can sort through the database for alloys which exhibit maximal values of the merit indices - the lightest, the most creep resistant, the most oxidation resistant, and the cheapest for example. Or alternatively, one can use the database to determine the relative trade-offs in performance which arise from different combination of properties. The six merit indicies are now described.

The first merit index is the creep-merit index. The overarching observation is that time- dependent deformation (i.e. creep) of a nickel-based superalloy occurs by dislocation creep with the initial activity being restricted to the γ phase. Thus, because the fraction of the γ' phase is large, dislocation segments rapidly become pinned at the γ/γ' interfaces. The rate-controlling step is then the escape of trapped configurations of dislocations from γ/γ' interfaces, and it is the dependence of this on local chemistry - in this case composition of the γ phase - which gives rise to a significant influence of alloy composition on creep properties.

A physically-based microstructure model can be invoked for the rate of accumulation of creep strain έ when loading is uniaxial and along the (OOl crystallographic direction. The equation set is

Ρ«, = ^€έ ( _∞ ι) (3) where p _m is the mobile dislocation density, φ _ρ is the volume fraction of the γ' phase, and ω is width of the matrix channels. The terms <J and T are the applied stress and temperature, respectively. The terms b and k are the Burgers vector and Boltzmann constant, respectively. The term K _CF - 1 + 2φ _ρ ¹¹³ / -φ _ρ ^ι ' ³ ) is a constraint factor, which accounts for the close proximity of the cuboidal particles in these alloys. Equation 3 describes the dislocation multiplication process which needs an estimate of the multiplication parameter C and the initial dislocation density. The term is the effective diffusivity controlling the climb processes at the particle/matrix interfaces.

Note that in the above, the composition dependence arises from the two terms φ _ρ and

-¾· . Thus, provided that the microstructural architecture is assumed constant (microstructural architecture is mostly controlled by heat treatment) so that φ _ρ is fixed, any dependence upon chemical composition arises through D _eS . For the purposes of the alloy design modelling described here, it turns out to be unnecessary to implement a full integration of Equations 2 and 3 for each prototype alloy composition. Instead, a first order merit index Μ _αβφ ^{1 'S} employed which needs to be maximised, which is given by c _reep (4) where , is the atomic fraction of solute / ^' in the γ phase and D _i is the appropriate interdiffusion coefficient.

The second merit index is a strength merit index. For high nickel-based superalloys, the vast majority of strength comes from the precipitate phase. Therefore, optimising alloy composition for maximal precipitate strengthening is a critical design consideration. From hardening theory a merit index for strength, M _strength , is proposed. The index considers the maximum possible precipitate strength - determined to be the point where the transition from weakly coupled to strongly coupled dislocation shearing occurs - which can be approximated using,

^strength = M. iy _APB 0j ^/2 /b (5)

Where M is the Taylor factor, y _APB is the anti-phase boundary (APB) energy, _ρ is the volume fraction of the γ' phase and b is the Burgers vector.

From Equation 5 it is apparent that fault energies in the γ' phase - for example, the antiphase boundary APB energy - have a significant influence on the deformation behaviour of nickel-based superalloys. Increasing the APB energy has been found to improve mechanical properties including, tensile strength and resistance to creep deformation. The APB energy was studied for a number of Ni-Al-X systems using density functional theory. From this work the effect of ternary elements on the APB energy of the γ' phase was calculated, linear superposition of the effect for each ternary addition was assumed when considering complex multicomponent systems, resulting in the following equation,

YAPB = I ⁹⁵ - 1.7x _Cr ^{~ 1 x} Mo + 4.6x _w + 27.1x _Ta + 21Ax _Nb + 15x _Ti (6) where, xcr, XMO, XW, χτα, m and xn represent the concentrations, in atomic percent, of chromium, molybdenum, tungsten, tantalum, niobium and titanium in the γ' phase, respectively. The composition of the γ' phase is determined from phase equilibrium calculations. The third merit index is density. The density, p, was calculated using a simple rule of mixtures and a correctional factor, where, , is the density for a given element and x, is the atomic fraction of the alloy element.

The fourh merit index is cost. In order to estimate the cost of each alloy a simple rule of mixtures was applied, where the weight fraction of the alloy element, x,, was multiplied by the current (2017) raw material cost for the alloying element, c,.

Cost = ∑i XiCi (8)

The estimates assume that processing costs are identical for all alloys, i.e. that the product yield is not affected by composition.

A fifth merit index is based upon rejection of candidate alloys on the basis of unsuitable microstructural architecture made on the basis of susceptibility to TCP phases. To do this use is made of the d-orbital energy levels of the alloying elements (referred as Md) to determine the total effective Md level according to d =∑ _i x _i Md _i (9) where the x, represents the mole fraction of the element i in the alloy. Higher values of Md are indicative of higher probability of TCP formation.

A sixth merit index is the gamma-prime solvus temperature. The gamma-prime solvus is defined as the temperature where the volume fraction of gamma-prime tends to zero. This is determined using thermodynamic calculations - as previously described above in the second step of the Alloys-by-Design method. The phase diagram and thermodynamic properties for a specific alloy composition is calculated and used to find the temperature at which this phase transition occurs.

The ABD method described above was used to isolate the inventive alloy composition. The design intent for this alloy was to deliver substantially high strengthen in combination with a reduction in alloy cost and an improvement in oxidation/corrosion resistance. The balance of properties for the new alloy make it more cost effective for the production of components for high temperature applications. In particular for use as a turbine disc application where the operating temperature for the component is 800°C or greater The material properties - determined using the ABD method - for the nominal compositions of the commonly applied/researched alloys used for powder metallurgy (PM) turbine disc applications, listed in Table 1, are listed in Table 3. The design of the new alloy was considered in relation to the predicted properties listed for these alloys. The rationale for the design of the new alloy is now described.

Table 3: Calculated phase fractions, misfit and merit indices made with the "Alloys-by-Design " software. Results for nickel-based superalloys listed in Table 1.

N18 0.55 -0.61% 2.81 1472 8.05 17.7 0.942 1194 3.1

Rene88DT 0.34 0.04% 2.45 1392 8.52 16.7 0.914 1108 2.7

RR1000 0.41 -0.07% 2.77 1474 8.31 21.3 0.921 1139 3.1

ME3 0.48 -0.05% 3.20 1608 8.34 23.1 0.922 1163 3.5

AlloylO 0.52 0.01% 3.16 1616 8.43 21.0 0.911 1182 3.4

LSHR 0.51 -0.11% 3.50 1647 8.40 22.4 0.932 1171 3.5

US8,147,749 0.50 0.35% 3.32 1854 8.76 28.6 0.906 1141 4.1 19 0.41 0.01% 2.41 1541 8.47 16.4 0.913 1136 3.1

US8,613,810 0.52 0.14% 3.22 1818 8.56 25.2 0.921 1156 4.1

US2015/0192002 0.51 0.17% 2.90 1719 8.55 21.6 0.921 1152 3.7

Identified on Figure 1 is the calculated cost merit index and strength merit index for the nominal compositions of the commonly applied/researched alloys used for powder metallurgy (PM) turbine disc applications listed in Table 1. The calculated cost and strength merit index for many millions of alloys which fall within the compositional design space described in Table 2 are also shown (the shaded area). For a given alloy cost there is an upper limit to the maximum achievable strength, Figure 1. The dotted line on Figure 1 represents the maximum achievable strength for any given alloy cost. It can be seen that it is possible to achieve high levels of strength (which are equivalent to or better than the prior art) whilst also reducing the elemental cost of the alloy. The hatched area on Figure 1 depicts the targeted performance criteria for the present invention in terms of alloy cost and strength merit index. The isolation of alloy compositions which provide a combination of high strength and low cost is described in the following sections.

Figures 2-6 show the influence of the elements tantalum, cobalt and niobium on alloy cost. These elements have a cost which is substantially higher than the base element nickel and therefore strongly increase the price of the alloy (based on metal September 2017 metal prices). The cost of tantalum, cobalt and niobium normalised to the price of nickel are 11.93, 5.45 and 3.82, respectively. A minimum cobalt level of 3.8 wt.% is desirable to provide the alloy with sufficient creep strength (in terms of creep merit index), described later in relation to Figures 15-20. The cost target for the present invention is less than or equal to 14.7 $/kg; reducing the elemental cost by 10% in comparison to lowest cost alloy listed in Table 1 (Alloy N19, 16.4$/Kg). From Figures 2-6 a relationship between elemental additions (based upon niobium, cobalt and tantalum) and target alloy cost was derived. The relationship - which describes a 3- dimensional surface - was determined in the following way. A relationship between alloy cost and niobium and cobalt was determined from Figures 2-6 for the cost target of 14.7 $/kg and the preferred cost targets of 14.0 $/kg and 13.1 $/kg, this was determined from the contour lines for these values. The translation of the lines as a function of different tantalum contents was determined from Figures 2-6. From the process described it is determined that the additions of tantalum, cobalt and niobium adhere to the following equation fXcost) = W _m + 1.53 W _Co + 3JW _Ta where ./(cost) is a numerical value and Wm, Wco and W-\ _a are the weight percent of niobium, cobalt and tantalum in the alloy respectively. For example, in order to produce an alloy with a cost of 14.7 $/kg or less the numerical value for ./(cost) should be less than or equal to 14.7. Based on the minimum cobalt levels of the present invention (3.8 wt.%) and minimum niobium concentration (0.0 wt.%) the maximum tantalum content in the alloy should be 2.4 wt.%, see /(cost) or less. The maximum cobalt content should be limited to 9.6 wt.% in order to achieve the cost target (Figure 2). Preferably the numerical value for /(cost) should be less than or equal to 12.4 to achieve a cost of 14.0 $/kg or less reducing the elemental cost by 15% compared to the lowest cost alloy (Alloy N19, 16.4$/Kg), therefore based on the minimum cobalt concentration (3.8 wt.%) and minimum niobium concentration (0.0 wt.%) it is preferred that the maximum tantalum content in the alloy should be less than or equal to 1.8 wt.%. The maximum cobalt content is preferably limited to less than 8.1 wt.% for cost considerations (Figure 2). Most preferable the numerical value for (cost) should be less than or equal to 9.5 to achieve a cost of 13.1 $/kg or less reducing the elemental cost by 20% compared to the lowest cost alloy (Alloy 19, 16.4$/Kg) therefore based on the minimum cobalt concentration (3.8 wt.%) and minimum niobium concentration (0.0 wt.%)it is preferred that the maximum tantalum content in the alloy should be less than or equal to 1.0 wt.%. To achieve a cost of 13.1 $/kg or less it is most preferable that the maximum cobalt content is preferably limited to 6.0 wt.% (Figure 2).

In an embodiment of the invention is desirable to include iron in substitution for nickel content. This has the benefit of reducing alloy cost and increasing the ability for the alloy to be recycled. Additions of iron may result in increased microstructural instability. Limiting iron additions to a level of 4.0 wt.% produces a good balance of low cost, improved recyclability and microstructural stability, more preferably a range between 1.0 wt.% and 2.0 wt.% is desirable.

In combination with a low cost the present invention relates to a high strength alloy. The target for the invention is to have a strength merit index of 1700 MPa or higher. This target for strength merit index provides the alloy with a maximum operating temperature equivalent to the highest strength alloys listed in Table 1 (three highest strength alloys ranging between 1719 -1854MPa), but at substantially reduced elemental cost (between 30-45% compared to alloys with strength merit index greater than ΠΟΟΜΡα). Thus, a substantial improvement in the combination of strength and cost is achieved.

The main alloying additions used to increase the strength merit index are the gamma- prime (γ') forming elements, aluminium, titanium, niobium and tantalum. Figures 7-11 show the influence of the elements aluminium, titanium and niobium on strength for alloys which meet the cost target (note tantalum containing alloys are not included in results, Figures 7-11). The addition of titanium, niobium and tantalum have been given a factor to convert the weight percent addition to an "aluminium equivalent". This allows for direct comparison of the influence of elements which have very different densities. For example titanium has a density of 4.5g/cm ³ compared to aluminium with density of 2.7g/cm ³ ' thus a factor of 0.6 is applied (i.e. 2.7/4.5=0.6). Similar to titanium, a constant is added to convert the elemental additions of niobium (8.57 g/cm ³ ) and tantalum (16.4 g/cm ³ ) to an "aluminium equivalent", thus, niobium and tantalum have correctional factors (determined from their density relative to aluminium) of 0.3 and 0.15 respectively. From Figures 7-11 a relationship between elemental additions (based upon titanium, aluminium and niobium) and target alloy strength was derived. The relationship - which describes a 3 -dimensional surface - was determined in the following way. A relationship between alloy strength and aluminium and titanium was determined from Figures 7-11 for the strength target of 1700MPa, this was determined for the contour lines for these values. The translation of the line as a function of different niobium contents was determined from Figures 7-1 1. From the process described it is determined that additions of aluminium, titanium and niobium adhere to the following equation f (strength) = 0.6W _Tl + 0.31W _Nb + 0.12W _M where strength) is a numerical value and Wn, Wm and WAX are the weight percent of titanium, niobium, aluminium in the alloy respectively. In order to produce an alloy with a strength merit index greater than or equal to 1700 MPa the numerical value for /(strength) should be greater than or equal to 3.7.

Tantalum is an optional element present in an amount of 2.4 wt% or less. In an embodiment of the invention tantalum can be added up to a level 2.4 wt.% in atomic substitution for titanium and niobium, (that is, for every atom of tantalum added the maximum of allowable level of one or both of titanium and niobium is decreased such that the maximum atomic percent of titanium plus tantalum plus niobium is equal to the atomic percent equivalent of the sum of the maximum allowable levels of titanium and niobium). .

Additions of aluminium, titanium, niobium and tantalum adhere to the following equation (note a factor of 0.27 is applied for tantalum instead of 0.15 determined based its influence on γ' formation because it has a significant influence on APB energy strengthening, see Equation 6. From equation 6 the APB strengthening factor for Tantalum is 27.1, the factor applied to titanium is 15, therefore tantalum is 1.8 times more potent in strengthening. This factor is multiplied by the aluminium equivalent of tantalum (0.15), thus a factor of 0.27 is derived (i.e. 0.15 x 1.80 = 0.27).

/ (strength) = 0.6W _Ti + 0.31W _Nb + 0.27W _Ta + 0.12W _M

In order to produce an alloy with a strength merit index greater than or equal to 1700 MPa the numerical value for/(strength) should be greater than or equal to 3.7. For even greater strength the numerical value for (strength) should preferably be greater than or equal to 4.0 or even 4.1.

In order to achieve a good balance between strength and ability to process the alloy it is required that the solvus temperature of the y'phase is less than 1180°C. A γ' solvus of less than 1180°C is preferred as this allows for heat treatment above the γ' solvus whilst reducing the susceptibility of the alloy to cracking on cool down from above the γ' solvus temperature. Heat-treatment above the γ' solvus temperature is desirable as this enables the growth of coarse grains which improving resistance to dwell fatigue, this damage mechanism is often a life limiting factor in this class of alloys. From Figure 12 a relationship between alloy γ' solvus and aluminium and the sum of elements titanium and niobium (according to the relationship 0.6WTi+0.31 WNt>) was determined. Additions of aluminium, titanium and niobium adhere to the following equation f (solvus) = 0.6W _Tl + 0.31W _Nb + 0.9W _M where /(solvus) is a numerical value. In an embodiment of the invention tantalum can be added up to a level 2.4 wt.% in substitution for titanium and niobium. The function solvus) is then: f (solvus) = 0.6W _Ti + 0.31W _Nb + 0.15W _Ta + 0.9W _M

In order to produce an alloy with a solvus of 1 180°C or less the value for /(solvus) should be less than 6.3, preferably the numerical value should be less than 5.9 to produce an alloy with a solvus less than 1 170°C which will improve the ability to process the alloy.

Based upon the equation which describes/(strength) and the need to have a y'solvus of 1 180°C or less the maximum aluminium content is limited to 3.3 wt.% or less (Figure 12), preferably to achieve γ' solvus of less than 1170°C aluminium content is limited to 3.0 wt.% or less (Figure 12). The ratio of the elements according to (0.6Wn+03 l Wm)/WM≤1.3 [is desired (see D.J. Crudden, N. Warnken, A. Mottura, and R.C. Reed. Modelling of the influence of alloy composition on flow stress in high-strength nickel-based superalloys. Acta Materialia, 7:356-370, 2014) as this is required to maintain stability of the y'phase, at values beyond this the formation of unwanted eta phase (Ni ₃ Ti) and delta phase (Ni ₃ Nb) are likely to occur, these phases reduce alloy ductility and fatigue resistance. Niobium must also be limited to less than 4.0 wt.% as high concentrations of niobium stabilise the niobium rich delta phase Ni ₃ Nb. Because of its deleterious effect on stability, niobium need not be present in the alloy. In an embodiment of the invention tantalum can be added up to a level 2.4 wt.% in substitution for titanium and niobium, in this embodiment the ratio of the elements should adhere to the following relationship (0.6#¾+0.31 W b+0.15 Wr _a WM <1.3

Based upon the maximum content of elements niobium (4 wt.%), aluminium (3.3 wt.%) and tantalum (2.4 wt.%) the minimum level of titanium required is 2.3 wt.% (see equation for /(strength)), preferably titanium is greater than 2.6 wt.% to allow for the preferred tantalum content of 1.8wt.%, more preferably the titanium content is greater than 3.0 wt.% to allow for the more preferable tantalum content (1.0%), most preferably titanium is greater than 3.5 wt.% so that the alloy can be substantially free from tantalum, this provides an improved combination of cost and strength. Even more preferably titanium is greater than 4.1 wt.% as this allows the alloy to be substantially free from tantalum and use the preferred level of niobium (2.7 wt.% or less) which is described later in relation to creep merit index and alloy cost . Based upon the maximum solvus temperature (1180°C) the sum of the elements according to the relationship 0.6W¾+0.31 iFNb+0.15WTa must be less than 3.33, based upon maximum aluminium content (3.3 wt.%)) and minimum niobium and tantalum contents (0.0 wt.%), see equation for/(solvus). Therefore the maximum allowable content for titanium is 5.6 wt.% (i.e. 3.33/0.6=5.6). The aluminium content should be greater than 2.6 wt.% to ensure that a desirable combination of strength, γ' solvus and stability of the γ' phase is achieved (i.e. (0.6 Wn+0 1

<1.3) (Figure 12). More preferably the aluminium content is limited to 3 wt.% to reduce γ' solvus (Figure 12).

The additions of aluminium, niobium and titanium influence the fraction of the precipitate hardening γ' phase. Based on the requirement for a strength merit index to be greater than 1700MPa - determined by the function for /(strength) - and the ratio of the elements (0.6Η¾+0.31 Wm)/W _A \ <1.3 the γ' fraction in the alloy should range between 51 % and 58% at a temperature of 850°C (Figure 13). The gamma-prime volume fraction is measured experimentally by the following procedure. After an aging heat treatment at 850°C a section is taken through the material and polished using conventional/standard metallurgical preparation techniques for scanning electron microscopy. Once prepared the gamma/gamma-prime microstructure should be observable in a scanning electron microscope, particles of diameter 30μηι or lower should be observable. A 10 of images are taken which provide a statistically representative dataset, the images should cover an area of at least 1mm ² . The 2-dimensional images which reveal the gamma/gamma-prime microstructure should be processed to identify the gamma-prime phase, the area fraction of the gamma-prime phase should be measured. The area fraction of the phase is taken to be the volume fraction of gamma-prime and should lie between 51-58%. From Figure 13 a relationship between alloy γ' fraction and aluminium and the sum of elements titanium.and niobium (according to the relationship 0.6WT _I +0.31 WNI _> ) was determined. To achieve an alloy with a γ' fraction between 51% and 58% the additions of aluminium, titanium and niobium adhere to the following equation

/( ) = 0.6W _TL + Q31W _NB + 0.8W _M where ί(γ') is a numerical value which ranges between 5.4 and 6.0.

In an embodiment of the invention tantalum can be added up to a level 2.4 wt.% in substitution for titanium and niobium, additions of aluminium, titanium, niobium and tantalum preferably adheres to the following equation

5.4 < 0.6W _Ti + 031W _Nb + 0.15W _Ta + 0.8W _M ≤ 6.0

Additions of tantalum are made to increase the strength of the alloy - in terms of strength merit index - by increasing γ' volume fraction and also by increasing the anti-phase boundary energy. Additions of tungsten to the alloy improve creep strength in the alloy - in terms of creep merit index - due to it very slow diffusion in the gamma matrix, strengthening is also increase by solid solution strengthening the gamma matrix phase. However, of the alloying elements listed in the design space in Table 2 the elements tungsten and tantalum are the heaviest, with a density substantially greater than nickel. Therefore their contribution to strength must be balanced against their negative impact upon alloy density. The density target for the present invention is less than 8.4 g/cm ³ , making it lighter than other alloys with a strength merit index of greater than 1700MPa (Table 3). Figure 14 shows the effect of the elements tantalum and tungsten on alloy density. The addition of tungsten is less than 6.2 wt.% such that a density of under 8.4 g/cm ³ is achieved. From Figure 14 it is determined that the additions of the elements tantalum and tungsten adhere to the following equation fidesnity) = W _W + Ta where_ (density) is a numerical value which is under 6.2 in order to attain a density of 8.4 g/cm ³ or less. Preferably the numerical value for^density) is less than 4.6 as this produces and alloy with a density of 8.3 g/cm ³ or less, even more preferably the numerical value for (density) is less than 2.9 as this produces and alloy with a density of less than 8.2 g/cm ³ . Thus it is preferable that tungsten is less than or equal to 4.6 wt.%, more preferably less than 2.9 wt.%. A target for creep merit index of 2.4 x 10 ^"15 m ^"2 s or greater is required for the alloy, to provide creep strength equivalent to AlloyN19 at areduced cost. The target density of 8.4g/cm ³ or less means tungsten is limited to less than 6.2 wt.% (Figure 14). The role of tungsten, cobalt and molybdenum is demonstrated in Figures 15-20. From Figures 15-20 a relationship between elemental additions (based upon cobalt, molybdenum and tungsten) and target alloy creep merit index was derived. The relationship - which describes a 3 -dimensional surface - was determined in the following way. A relationship between alloy creep merit index and cobalt and tungsten was determined from Figures 15-20 for the creep merit index target of 2.4 x 10 ^{" 15} m ^"2 s , this was determined for the contour lines for these values. The translation of the line as a function of different molybdenum contents was determined from Figures 15-20. From the process described to increase creep resistance - in terms of creep merit index - additions of tungsten, cobalt and molybdenum adhere to the follow equation f (creep) = W _W + 0.96W _Co + 1.2W _Mo where (creep) is a numerical value which should be greater than 16.0 in order to achieve a creep merit index of greater than 2.4 x 10 ^"15 m ^"2 s. Therefore based upon the maximum levels of tungsten (6.2 wt.%) and cobalt (9.6 wt.%) it is necessary to have at least 0.6 wt.% molybdenum to attain sufficient creep resistance. Preferably tungsten is limited to 4.6 wt.% to lower alloy density therefore it is preferable to increase molybdenum to 2.2 wt.% or greater. Even more preferable tungsten is limited to 4.6 wt.% and cobalt is less than 8.1 wt.% to reduce both alloy density and cost, therefore it is preferable that that molybdenum is greater than 3.6 wt.%). The maximum molybdenum content limited to less than 5.1 wt.% due to stability requirements, explain later with reference to Figure 21.

A minimum tungsten content of 0.7 wt.% is required at maximum cobalt of 9.6 wt.% to achieve the minimum creep merit index of 2.4 x 10 ^"15 m ^"2 s, Figure 20. Preferably molybdenum is limited to 4.1% for improved alloy stability (see below), thus a tungsten content of greater than 1.9 wt.% is preferred as this results in a better balance of creep resistance and alloy stability. Based upon the maximum levels for tungsten (6.2 wt.%) and molybdenum (5.1 wt.%)), a minimum concertation of 3.8 wt.% cobalt is desired to achieve a good level of creep resistance (Figure 20). More preferably tungsten is limited to 4.6 wt.%, resulting in a preferred cobalt level of 5.5 wt.% or more, even more preferably the tungsten content is limited to 2.9 wt.%) therefore a cobalt content of 8.0 wt.% or greater is preferable (Figure 20) for balance of creep resistance and alloy stability. Based upon the preferred cobalt content of 8.0 wt.% or greater for an improved balance of creep resistance and density it is preferred that niobium is limited to less than 2.7 wt.%> (Figure 2).

Figure 21 describes the effect of chromium content and the sum of the elements molybdenum and tungsten (according to the relationship Mo+0.5W) on the stability number. A higher stability number results in an alloy which is more prone to TCP phase formation. Limiting or stopping the precipitation of TCP phase formation is beneficial as these phases lead to deterioration in material properties. The correctional factor of 0.5 is applied to tungsten as it has a density approximately twice that of molybdenum, this factor accounts for the substantial difference in density of the elements. A chromium level of greater than 9.5 wt.% is desirable in order to achieve a good level of oxidation resistance. A stability target of less than 0.91 in order to ensure microstructural stability and avoid TCP formation. Therefore a molybdenum content of less than 5.1 wt.% (md<=0.91) is desired, preferably 4.1 wt.% or less (md<=0.90), Figure 21. Based on creep requirement as determined by /(creep), the minimum sum of molybdenum and tungsten according to the relation (Mo+0.5W) is 3.7 wt.% (maximum tungsten content of 6.1 wt.% and minimum molybdenum content of 0.6 wt.%). Therefore maximum chromium level must be 14.4 wt.%. Preferably the chromium content is greater than 10.0% as this gives an improvement in oxidation resistance. More preferably the chromium content is greater than 1 1.5 wt.% or even 12.0 wt.% as this will give an improved oxidation resistance (higher Cr content) than other alloys which have a strength merit index of greater than 1700 MPa. From Figure 21 it is determined that the additions of the elements molybdenum tungsten and chromium adhere to the following equation /(stability) = W _Mo + 0.514V + 0.8M^ _r where /(stability) is a numerical value which is under 15.5 in order to maintain alloy stability.

Additions of carbon, boron and zirconium are required in order to provide strength to grain boundaries. This is particularly beneficial for the creep and fatigue properties of the alloy. Carbon is added to act as a grain boundary pinning particle, this is necessary when heat treatment is conducted above the gamma-prime solvus temperature to inhibit excessive grain growth. The carbon concentrations should range between 0.01 wt.% and 0.1 wt.%. Preferably, the levels of carbon are between 0.2 and 0.06, this range provides a better distribution of carbide phases for controlling alloy microstructure, in particular grain size.

The boron concentration should range between 0.001 and 0.1 wt.%. The addition of boron can improve creep ductility and grain boundary strength through the formation of boride phases. Preferably the boron content in the alloy is between 0.01 and 0.05 wt.%, as this provides a desirable level of the boride phase.

The zirconium concentrations should range between 0.001 wt.% and 0.3 wt.%, preferably between 0.02 and 0.1 wt.%. Zirconium plays a role in guttering unwanted impurities in the ally, for example, oxygen and sulphur. These impurities may lead to embrittlement of the alloy particularly due to grain boundary embrittlement. It is beneficial that when the alloy is produced, it is substantially free from incidental impurities. These impurities may include the elements sulphur (S), manganese (Mn) and copper (Cu). The element sulphur should remain below 0.003 wt.% (30 PPM in terms of mass). Manganese is an incidental impurity which is limited to 0.25 wt.%, preferably this limited to less than 0.1 wt.%. Copper (Cu) is an incidental impurity which is preferably limited to 0.5 wt.%. The presence of Sulphur above 0.003 wt.%, can lead to embrittlement of the alloy and sulphur also segregates to alloy/oxide interfaces formed during oxidation, preferably sulphur levels of less than less than 0.001 wt.%. Vanadium is an incidental impurity, vanadium negatively influences the oxidation behaviour of the alloy and is which is preferably limited to 0.5 wt.%, preferably less than 0.3 wt.% and most preferably this limited to less than 0.1 wt.%. This segregation may lead to increased spallation of protective oxide scales. If the concentrations of these incidental impurities exceed the specified levels, issues surrounding product yield and deterioration of the material properties of the alloy is expected.

Additions of hafnium (Hf) of up to 0.5wt.%, are beneficial for tying up incidental impurities in the alloy and also for providing strength. Hafnium is a strong carbide former it can provide additional grain boundary strengthening. More preferably hafnium is limited to 0.2wt.%, more preferably less than 0.1 wt.% as the elemental cost is significant, additions have a negative impact on alloy cost.

Additions of the so called 'reactive-elements', Yttrium(Y), Lanthanum (La) and Cerium (Ce) may be beneficial up to levels of 0.1 wt.% to improve the adhesion of protective oxide layers, such as Cr ₂ 0 ₃ . These reactive elements can 'mop-up' tramp elements, for example sulphur, which segregates to the alloy oxide interface weakening the bond between oxide and substrate leading to oxide spallation. Additions of Silicon (Si) up to 0.5 wt.%) may be beneficial, it has been shown that additions of silicon to nickel based superalloys at levels up to 0.5 wt.% are beneficial for oxidation properties. In particular silicon segregates to the alloy/oxide interface and improves cohesion of the oxide to the substrate. This reduces spallation of the oxide, hence, improving oxidation resistance.

Based upon the description of the invention presented in this section the broad range for the invention is listed in Table 4. A preferable range is also given in Table 4. 2018/052973

23

Table 4: Compositional range in wt. % for the newly design alloy.

Alloy (wt.%) Cr Co Fe Mo W Al Ti Ta Nb

Min 9.5 3.8 0.0 0.6 0.7 2.6 2.3 0.0 0.0

Max 14.4 9.6 4.0 5.1 6.2 3.3 5.6 2.4 4.0

Preferable Min 10.0 5.5 1.0 2.2 1.9 2.6 2.6 0.0 0.0

Preferable Max 14.4 8.1 2.0 4.1 4.6 3.0 5.6 1.8 2.7

Most Preferable Min 1 1.5 5.5 1.0 3.6 1.9 2.6 4.1 0.0 0.0

Most Preferable Max 14.4 6.0 2.0 4.1 2.9 3.0 5.6 1.0 2.7

The following Section describes example compositions for the present invention. The calculated properties for these new alloys are listed. The rationale for the design of these alloys is now described.

Examples of the Invention

An example of the composition invention (Example LCPM-1) is listed in Table 5. The alloy listed has a number of benefits in comparison to the high strength alloys listed in Table 1 with a strength merit index greater than or equal to 1700MPa (Table 3). In particular, the cost is substantially lower ($14.5 $/kg) in comparison to these alloys with cost between (21.6-28.6 $/kg), providing at least a 32% reduction in elemental cost whilst maintaining an equivalent strength. The LCPM alloy is also of lower density 8.3 g/cm ³ in comparison to these alloys with density ranging between 8.5-8.7 g/cm ³ . The LCPM alloy has a similar chromium content (11.8 wt.%) equivalent to the these alloys which range between 10.0-13.4 wt.%. In Table 5 the ratio of the elements according to the relationship

(0.6 J _T i+0.31 ^Nb+0.15WTa)/^Ai has been modified between 1.13-1.28. It is seen that by doing this the strength of the alloy is increased further (Table 6).

Table 5: Example compositions where the ratio of (0.6Wn+0.3 lWm+0.1 ^' 5Wr _a )/WAi has been modified between 1.13-1.28

Alloy (wt.%) Cr Co Fe Mo W Al Ti Ta Nb C B Zr

LCPM-1 11.8 8.0 0.0 4.0 4.0 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-2 11.8 8.0 0.0 4.0 4.0 2.9 4.6 0.0 2.4 0.050 0.030 0.060

LCPM-3 11.8 8.0 0.0 4.0 4.0 2.8 4.7 0.0 2.5 0.050 0.030 0.060 Table 6: Calculated phase fractions and merit indices made with the "Alloys-by-Design " software. Results for compositions where the ratio of (0.6 J^Ti+0.31 i-Fkb+0.15WT ₃ )/WAI has been modified between 1.13-1.28 w example alloy LCPM-1 (T ble 5).

γ/γ- Creep Strength 0.6IV _Ti -t- 0.31MV

Alloy (wt.%) Y ¹ Misfit Merit Index Merit Index Density Cost Md _Y γ' Solvus + 0.27^ + 0.12^

(m ^' ¾ x 10 ^"

(%) (Mpa) (g/cm ³ ) ($/kg) (°C)

1 ⁵ )

LCPM-1 0.52 -0.10% 2.4 1710 8.3 14.5 0.92 1172 3.7

LCPM-2 0.52 -0.07% 2.4 1739 8.3 14.5 0.92 1171 3.8

LCPM-3 0.52 -0.05% 2.4 1768 8.3 14.6 0.92 1170 3.9

5 In Table 7 the composition of example alloy LCPM-1 is modified in order to achieve and alloy with a lower cost. Cobalt is removed for the alloy to reduce alloy cost. As cobalt provides creep resistance to the alloy substitution of cobalt using molybdenum and tungsten is required, according to the relationship defined for (creep). To manage alloy stability, according to the relationship for (stability) chromium is reduced as molybdenum and tungsten0 levels are increased. The alloys LCPM 6-8 are beneficial when a there is a stronger preference for reduced alloy cost and a less strong preference for lower alloy density (Table 8) and better alloy corrosion resistance (i.e. chromium levels are reduced)

Table 7: Example compositions where the chromium, cobalt, molybdenum and tungsten content have been modified in example alloy LCPM-1 to maintain alloy creep resistance at a lower cost.

Alloy (wt.%) Cr Co Fe Mo W Al Ti Ta Nb C B Zr

LCPM-1 11.8 8.0 0.0 4.0 4.0 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-6 11.0 7.0 0.0 4.4 4.5 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-7 10.2 6.0 0.0 4.8 5.0 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-8 9.5 5.0 0.0 5.2 5.4 3.0 4.5 0.0 2.2 0.050 0.030 0.060

5 Table 8: Calculated phase fractions and merit indices made with the "Alloys-by-Design " software.

Results for compositions the chromium, cobalt, molybdenum and tungsten have been modified to maintain alloy creep resistance at a lower cost in example alloy LCPM-1 (Table 7).

Creep 0.6W _T1 + 0.3111^,, γ/γ' Merit Strength + 0. ZlW _Ta

Alloy (wt.%) Misfit Index Merit Index Density Cost Md _y γ' Solvus + 0.12W _A ,

(rrr ² s x

(%) (Mpa) (g/cm ³ )

10- ¹⁵ ) ($/kg) (°C)

LCPM-1 0.52 -0.10% 2.4 1710 8.3 14.5 0.91 1172 3.7

LCPM-6 0.52 -0.17% 2.4 1712 8.3 14.1 0.91 1171 3.7

LCPM-7 0.53 -0.24% 2.4 1712 8.4 13.6 0.91 1171 3.7

LCPM-8 0.53 -0.33% 2.4 1713 8.4 13.2 0.91 1171 3.7

In Table 9 example alloy LCPM-7 is modified to include the optional element tantalum. It is demonstrated that tantalum can be substituted for titanium (LCPM 13-14) and niobium (LCPM 15-18). It is seen that when tantalum is added in atomic substitution for titanium or niobium there is an increase in alloy cost. The addition of tantalum as demonstrated by the examples is beneficial when an increase in alloy strength is preferred over a reduction of alloy cost.

Table 9: Example compositions where the niobium, titanium and tantalum content have been modified in example alloy LCPM-7.

Alloy (wt.%) Cr Co Fe Mo W Al Ti Ta Nb C B Zr

LCPM-7 10.2 6.0 0.0 4.8 5.0 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-13 10.2 6.0 0.0 4.8 5.0 3.0 4.4 0.4 2.2 0.050 0.030 0.060

LCPM-14 10.2 6.0 0.0 4.8 5.0 3.0 4.3 0.8 2.2 0.050 0.030 0.060

LCPM-15 10.2 6.0 0.0 4.8 5.0 3.0 4.5 0.2 2.1 0.050 0.030 0.060

LCPM-16 10.2 6.0 0.0 4.8 5.0 3.0 4.5 0.4 2.0 0.050 0.030 0.060

LCPM-17 10.2 6.0 0.0 4.8 5.0 3.0 4.5 0.6 1.9 0.050 0.030 0.060

LCPM-18 10.2 6.0 0.0 4.8 5.0 3.0 4.5 0.8 1.8 0.050 0.030 0.060

Table 10: Calculated phase fractions and merit indices made with the "Alloys-by-Design " software. Results for compositions where niobium, titanium and tantalum content have been modified in example alloy LCPM-7 (Table 9).

Creep 0.6W _Ti + 0.31W _wi , γ/γ. Merit Strength + 0.27W _Ta

Alloy (wt.%) 2Wn

Y' Misfit Index Merit Index Density Cost Md _r γ' Solvus + 0. i.

(m' ² s x

(%) (Mpa) (g cm ³ )

10- ¹⁵ ) (Meg) (°C)

LCPM-7 0.53 -0.24% 2.4 1712 8.4 13.6 0.91 1171 3.7

LCPM-13 0.53 -0.25% 2.4 1731 8.4 14.1 0.91 1170 3.8

LCPM-14 0.53 -0.26% 2.4 1750 8.4 14.6 0.91 1168 3.9

LCPM-15 0.53 -0.25% 2.4 1724 8.4 13.9 0.91 1171 3.8

LCPM-16 0.53 -0.26% 2.4 1735 8.4 14.1 0.91 1171 3.8

LCPM-17 0.54 -0.26% 2.4 1747 8.4 14.3 0.91 1170 3.9

LCPM-18 0.54 -0.27% 2.5 1759 8.4 14.6 0.91 1170 3.9 In Table 1 1 example alloy LCPM-1 the balance of molybdenum and tungsten in the alloy is modified. It is demonstrated that if molybdenum is reduced and tungsten is increased according to the relationship for /(creep) it is possible further to increase chromium levels in the alloys according to the relationship /(stability). The examples demonstrate when the ratio of elements tungsten to molybdenum is increased when there is an improvement in corrosion resistance due to increased chromium levels the trade-off associated with this is an increase in alloy density (Table 8).

Table 11: Example compositions where the chromium, molybdenum and tungsten content have been modified in example alloy LCPM-1 to adjust the balance between creep resistance and corrosion resistance.

Alloy (wt.%) Cr Co Fe Mo W Al Ti Ta Nb C B Zr LCPM-1 11.8 8.0 0.0 4.0 4.0 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-19 11.9 8.0 0.0 3.8 4.2 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-20 12.0 8.0 0.0 3.6 4.5 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-21 12.1 8.0 0.0 3.4 4.7 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-22 12.2 8.0 0.0 3.2 5.0 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-23 12.3 8.0 0.0 3.0 5.2 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-24 12.4 8.0 0.0 2.8 5.4 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-25 12.5 8.0 0.0 2.6 5.7 3.0 4.5 0.0 2.2 0.050 0.030 0.060

LCPM-26 12.6 8.0 0.0 2.4 5.9 3.0 4.5 0.0 2.2 0.050 0.030 0.060

Table 12: Calculated phase fractions and merit indices made with the "Alloys-by-Design " software. Results for compositions where the chromium, molybdenum and tungsten content have been modified in example alloy LCPM-1 to adjust the balance between creep resistance and corrosion resistance

(Table 11).

0.6W _Ti + .31W _Nb γ/γ' Creep Merit Strength Merit + 0.27W _Ta

Alloy (wt.%) γ' Misfit Index Index Density Cost Md _y γ' Solvus + 0.12W _M

(%) (m-¾x 10- ¹⁵ ) ( pa) (g/cm ³ ) ($/kg) (°C)

LCPM-1 0.52 -0.10% 2.4 1710 8.3 14.5 0.91 1172 3.7

LCPM-19 0.52 -0.09% 2.4 1712 8.3 14.5 0.91 1172 3.7

LCPM-20 0.52 -0.07% 2.5 1714 8.3 14.5 0.91 1172 3.7

LCPM-21 0.52 -0.06% 2.5 1717 8.3 14.5 0.91 1172 3.7

LCPM-22 0.53 -0.05% 2.5 1719 8.3 14.4 0.91 1172 3.7

LCPM-23 0.53 -0.04% 2.5 1721 8.3 14.4 0.91 1172 3.7

LCPM-24 0.53 -0.03% 2.5 1723 8.3 14.4 0.91 1172 3.7

LCPM-25 0.53 -0.02% 2.5 1725 8.4 14.4 0.91 1172 3.7

LCPM-26 0.53 -0.01% 2.6 1727 8.4 14.4 0.91 1172 3.7