WO2020254832A1

WO2020254832A1 - A nickel-based alloy

Info

Publication number: WO2020254832A1
Application number: PCT/GB2020/051501
Authority: WO
Inventors: Daniel BARBA CANCHO; Roger Charles REED; Enrique ALABORT MARTINEZ
Original assignee: Alloyed Limited
Priority date: 2019-06-21
Filing date: 2020-06-22
Publication date: 2020-12-24
Also published as: GB201908925D0; GB2584905B; GB2584905A; EP3987071A1

Abstract

A nickel-based alloy composition consisting, in weight percent, of: 3.0 to 6.5% chromium, 4.0 to 12.0% cobalt 1.5 to 7.5% tungsten, 0.0 to 0.5% molybdenum, 3.5 to 8.5% rhenium, 2.0 to 4.1% ruthenium, 5.3 to 6.8% aluminium, 6.0 to 9.7% tantalum, 0.0 to 0.5% hafnium, 0.0 to 0.5% niobium, 0.0 to 0.5% titanium, 0.0 to 0.5% vanadium, 0.0 to 1.0 platinum, 0.0 to 1.0 palladium, 0.0 to 1.0 iridium,0.0 to 0.1% silicon, 0.0 to 0.1% yttrium, 0.0 to 0.1% lanthanum, 0.0 to 0.1% cerium, 0.0 to 0.1% magnesium, 0.0 to 0.003% sulphur, 0.0 to 0.05% manganese, 0.0 to 0.05% zirconium, 0.0 to 0.005% boron, 0.0 to 0.01% carbon, 0.0 to 0.5% copper, 0.0 to 4.0% iron, the balance being nickel and incidental impurities.

Description

A nickel-based alloy

The present invention relates to a nickel-based single crystal superalloy composition designed for high performance jet propulsion applications. The alloy - a fourth generation single crystal nickel-based superalloy - exhibits a combination of creep resistance and oxidation resistance which is comparable to or better than equivalent grades of alloy. The density, cost, yield strength, processing and long-term stability of the alloy have also been considered in the design of the new alloy.

Examples of typical compositions of fourth generation nickel-based single crystal superalloys are listed in Table 1. These alloys may be used for the manufacture of rotating/stationary turbine blades used in gas turbine engines. Figure 1 shows schematically the trade off between creep resistance and strength and oxidation resistance made in some prior art alloys. Table 1: Nominal composition in wt. % of commercially used fourth generation single crystal turbine blade alloys.

Alloy A1 Cr Co Mo Re Ru W Ti Ta Hf

PW1497 5.6 2.0 16.5 2.0 6.0 3.0 6.0 OO 8 02

TMS-162 5.8 2.9 5.8 3.9 4.9 6.0 5.8 0.0 5.6 0.1

TMS-138 5.8 3.2 5.8 2.8 5.8 3.6 5.6 0.0 5.6 0.0

TMS-138A 5.7 3.2 5.8 2.8 5.8 3.6 5.6 0.0 5.6 0.0

MC-NG 6.0 4.0 0 1 4 4 5 0.5 5.0 0.1

MX4 5.55 2.0 16.5 2 6.0 3 6 0.0 8.3 0.2

ABD2 6.4 4.0 9 0 5.6 2.6 7.4 0.0 5.6 0.0

TMS-196 5.7 4.4 5.3 2.4 6.1 4.8 4.8 0.0 5.3 0.1

ABD2 is an alloy composition postulated in R. C. Reed et al,“Isolation and testing of new single crystal superalloys using alloys-by-design method” Materials Science and Engineering : A, Vol 667, 14 June 2016, pp 261-278.

It is an aim of the invention is to provide an alloy which has similar or improved high temperature behaviour in comparison to the fourth generation alloys listed in Table 1. The present invention provides a nickel-based alloy composition consisting, in weight percent, of: 3.0 to 6.5% chromium, 4.0 to 12.0% cobalt 1.5 to 7.5% tungsten, 0.0 to 0.5% molybdenum, 3.5 to 8.5% rhenium, 2.0 to 4.1% ruthenium, 5.3 to 6.8% aluminium, 6.0 to 9.7% tantalum, 0.0 to 0.5% hafnium, 0.0 to 0.5% niobium, 0.0 to 0.5% titanium, 0.0 to 0.5% vanadium, 0.0 to 1.0 platinum, 0.0 to 1.0 palladium, 0.0 to 1.0 iridium, 0.0 to 0.1% silicon, 0.0 to 0.1% yttrium, 0.0 to 0.1% lanthanum, 0.0 to 0.1% cerium, 0.0 to 0.1% magnesium, 0.0 to 0.003% sulphur, 0.0 to 0.05% manganese, 0.0 to 0.05% zirconium, 0.0 to 0.005% boron, 0.0 to 0.01% carbon, 0.0 to 0.5% copper, 0.0 to 4.0% iron, the balance being nickel and incidental impurities. This composition provides a good balance of properties, particularly between creep and oxidation resistance as well as cost, density, manufacturability and yield stress.

In an embodiment, the nickel-based alloy composition consists, in weight percent, of between 4.0 and 5.0% chromium. Such an alloy is particularly resistant to TCP formation and presents a high strength at mid temperatures whilst still having good oxidation & corrosion resistance at higher temperatures.

In an embodiment, the nickel-based alloy composition consists, in weight percent, of between 5.0% or 7.5% or 7.8% or 8.0% to 10.5 or 10.0% cobalt. Such an alloy has improved environmental resistance to oxidation combined with an increased resistance to deformation at mid temperatures and a limit in cost.

In an embodiment, the nickel-based alloy composition consists, in weight percent of, between 2.0 and 5.2 or 5.0% tungsten. This composition assures a good creep performance, high strength at low and mid temperatures while still achieving a light and stable alloy.

In an embodiment, the nickel-based alloy composition consists, in weight percent, of between 5.5 and 6.6% aluminium. This composition achieves the correct g’ amount for high creep resistance and strength while keeping a reduced density alongside with an increased oxidation resistance.

In an embodiment, the nickel-based alloy composition consists, in weight percent, of between 6.5 and 8.0% tantalum. This provides the optimal microstructure range with a good creep resistance, ease of manufacture (based upon solutioning window) and high strength at mid and low temperatures and reduces the cost and density of the alloy further and provides a superior strength up to higher temperatures, while keeping a good the solutioning window.

In an embodiment, the nickel-based alloy composition consists, in weight percent, of 0.1% or more molybdenum. This is advantageous for improved creep resistance.

In an embodiment, the nickel-based alloy composition consists, in weight percent of, between 5.0 and 6.5% rhenium, providing a good balance of creep resistance, density, resistance to TCP formation and cost.

In an embodiment, the nickel-based alloy composition consists, in weight percent of, between 2.5% or 2.8% or 3.0% and 4.0% or 3.5% ruthenium. This composition provides a good balance of creep resistance and cost and oxidation resistance.

In an embodiment, the nickel-based alloy composition consists, in weight percent, of between 0.0 and 0.2% hafnium. This is optimum for tying up incidental impurities in the alloy, for example, carbon.

In an embodiment, the nickel-based alloy composition is such that the following equation is satisfied in which W_T3 and WAI are the weight percent of tantalum and aluminium in the alloy respectively 34 < W_T3 + 5.0 WAI £ 39. This is advantageous as it allows a suitable volume fraction g' to be present (60-70%).

In an embodiment, the nickel-based alloy composition is such that the following equation is satisfied in which W_¾ and Wc_r are the weight percent of tantalum and chromium in the alloy respectively 3.5 < Wc_r - W_¾ ; preferably 3.0 < Wc_r - W_¾. This is advantageous as it allows a suitable solutioning window for the alloy to allow for heat-treatment processes.

In an embodiment, the nickel-based alloy composition is such that the following equation is satisfied in which W_RU and Wit_e are the weight percent of ruthenium and rhenium in the alloy respectively 4.8 > W_RU + 0.22 W_Re; preferably 3.7 > W_RU + 0.22 W_Re. This is advantageous as it results in an alloy with a relatively low cost.

In an embodiment, the nickel-based alloy composition is such that the following equation is satisfied in which W_Re and Ww are the weight percent of rhenium and tungsten in the alloy respectively 1.09 Wite + Ww £ 13.6_; preferably 1.09 Wite + Ww £ 11.6. This is advantageous as it results in an alloy with a relatively low density.

In an embodiment, the nickel-based alloy composition is such that the following equation is satisfied in which Wit_e, W_MO and Ww are the weight percent of rhenium, molybdenum and tungsten in the alloy respectively 0.29WR_U + O.lWRe +0.026Ww ³ 1.81; preferably 0.29WR_U + O.lWRe +0.026Ww ³ 1.86. This is advantageous as it results in an alloy with a high creep resistance.

In an embodiment, the nickel-based alloy composition is such that the following equation is satisfied in which Wc_o, Wc_r, Ww and W_ta are the weight percent of cobalt, chromium, tungsten and tantalum in the alloy respectively 1.10(W_Co + W_Cr)— (W_Ta + W_w) < 2.15. This is advantageous as it results in an alloy with a mid-temperature strength.

In an embodiment, in the nickel-based alloy composition, the sum of the elements niobium, titanium and vanadium, in weight percent, is less than 1%, preferably 0.5% or less. This means that those elements do not have too much of a deleterious effect on environmental resistance of the alloy.

In an embodiment, a nickel -based superalloys following 0.46 WAI + W_T3 > 3.34. This results in a suitable yield strength for the preferred g’ fraction (60-70%).

In an embodiment, the combination of W and Co following the relation 0.63Wc_o - Ww > -3.6; preferably 0.63Wc_o - Ww > 0.83 to maintain a good oxidation resistance.

In an embodiment, the nickel-based alloy composition has between 60 and 70% volume fraction g'.

In an embodiment, a single crystal article is provided, formed of the nickel-based alloy composition of any of the previous embodiments.

In an embodiment, a turbine blade for a gas turbine engine is provided, formed of an alloy according to any of the previous embodiments. In an embodiment, a gas turbine engine comprising the turbine blade of the previous embodiment is provided.

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%.

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

Figure 1 shows the evolution of the predicted values of oxidation, creep strength and yield strength as a function of the superalloy generation and with arrows indicate the advantage of the current invention over a conventional 4^th generation superalloy (same creep strength but with a good oxidation resistance and yield strength)

Figure 2 shows schematically the link between chemical elements and each of the alloy properties. For each element, the properties defining the upper and lower compositional limits are indicated.

Figure 3 is a contour plot showing the effect of g' forming elements aluminium and tantalum on volume fraction of g' for alloys within the alloy design space, determined from phase equilibrium calculations conducted at 900°C.;

Figure 4 is a contour plot showing the effect rhenium and tungsten on density, for alloys with a volume fraction of g' between 60-70% at 900°C with tantalum between 5 - 9 wt.%;

Figure 5 is a contour plot showing the effect of rhenium and ruthenium content on raw elemental cost, for alloys with a volume fraction of g' between 60-70% at 900°C with tantalum between 5 - 9 wt.%;

Figures 6a-e are contour plots showing the effect of elements chromium and tungsten on microstructural stability, for alloys with a volume fraction of g' between 60-70% at 900°C with tantalum between 5-9 wt.% and between 1-3 wt.% ruthenium, which contain 4 wt.% rhenium, 5 wt.% rhenium, 6 wt.% rhenium, 7 wt.% rhenium, respectively; Figure 7 is a contour plot showing the effect of elements aluminium and tantalum on the solutioning window for alloys with a volume fraction of g' between 60-70% at 900°C;

Figure 8 is a contour plot showing the effect of elements tungsten and cobalt on the oxidation resistance for alloys with a volume fraction of g' between 60-70% at 900°C with tantalum between 5-9 wt.%, and rhenium between 4-6wt.%

Figure 9 is a contour plot showing the effect of the elements tungsten and tantalum on the yield strength at room temperature for alloys with a volume fraction of g' between 60-70% at 900°C with tantalum between 5-9 wt.%, and rhenium between 4-6wt.%.

Figure 10 is a contour plot showing the effect of the combined elements (Cobalt + Chromium) and (Tantalum + Tungsten) on the resistance to plastic deformation at mid temperatures (600-800 °C) for alloys with a volume fraction of g' between 60-70% at 900°C with ruthenium 2-4 wt.%, and rhenium between 4-6wt.%.

Figure 1 la-d are contour plots showing the effect of elements rhenium and tungsten on the creep resistance, for alloys with a volume fraction of g' between 60-70% at 900°C with tantalum between 5-9 wt.%, which contain, 0 wt.% ruthenium, 1 wt.% ruthenium, 2 wt.% ruthenium, 3 wt.% ruthenium, respectively;

Figure 12 indicates in a triangular plot the performance of each of 3 example baselines alloys of each generation of yield strength, mid temperature creep resistance and high temperature creep resistance and comparison with the performance of the example alloys proposed.

Figure 13 shows the correlation between the calculated LMP_Mid and the measured experimental ones. R=0.77

Figure 14 shows the correlation between the calculated LMPcreep and the measured experimental ones. R=0.9 In superalloys, generally additions of chromium (Cr) and aluminium (Al) are added to impart resistance to oxidation and sulphidisation, cobalt (Co) is added to improve resistance to sulphidisation. For high temperature creep resistance, molybdenum (Mo), tungsten (W), Co, rhenium (Re) and sometimes ruthenium (Ru) 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) and titanium (Ti) are introduced as these promote the formation of the precipitate hardening phase gamma- prime (g'). This precipitate phase is coherent with the face-centered cubic (FCC) matrix phase which is referred to as gamma (y). For mid temperature creep resistance, the ration between g' and g stabilisers is increased (Ta+W)/(Co+Cr).

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 combined with machine learning 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. A diagram of the process is detailed in Fig. 2.

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 and in Fig. 2. The connection between each property and the affecting elemental limit is presented in Fig. 3 as a summary of the invention process.

Table 2: Alloys design space in wt. % searched using the“Alloys-by-Design” method.

Cr Co Re W AΪ Ta Ru

Min TO TO TO TO TO TO TO

Max 10.0 16.0 10.0 11.0 8.0 12.0 8.0

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 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 single crystal superalloys which require superior resistance to creep deformation, the creep rupture life is maximised when the volume fraction of the precipitate hardening phase g' lies between 60%-70%.

Rejection of alloy on the basis of unsuitable microstructural architecture is also made from estimates of susceptibility to topologically close-packed (TCP) phases based on the effective valence number of the g phase (Md·,). The present calculations predict the formation of the deleterious TCP phases (sigma (s), P and mu (m)) using CALPHAD modelling.

Thus the model isolates all compositions in the design space which are calculated to result in (1) a volume fraction of g' of between 60 and 70.

In the fourth stage, merit indices are estimated for the remaining isolated alloy compositions in the dataset based on physical models combined with machine learning tools using an extensive experimental alloy performance database. Examples of these include: (2) density, (3) cost, (4) solutioning window, (5) microstructural stability, (6) oxidation resistance, (7) yield strength of the alloy at room temperature, (8) Larson Miller parameter (LMP) for mid temperature resistance and (9) LMP for high temperature creep (which describes an alloy’s creep resistance based solely on mean composition). These indexes are indicated in Fig. 2.

In the fifth stage, the calculated merit indices are compared with limits for the required requirements, these design constraints are considered to be the boundary conditions to the problem. All compositions which do not fulfil the boundary conditions are excluded and the preferred compositional ranges are delimited. This process is presented in Fig. 2.

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. In this patent a combination of both procedures is used. First, the critical microstructural and physical requirements are imposed (1-6). Then, for the suitable alloys, a multitarget optimisation on the performance of the alloy at room, mid and high temperatures (7-9) is performed to rank the alloys.

The example 8 merit indices (2-9) are now described.

Physical Indexes

Density

The first merit index is density (2). The density, p, was calculated using a simple rule of mixtures and a correctional factor, where, p_t is the density for a given element and x, is the atomic fraction of the alloy element.

p = 1.05 Ei XiPi] (1)

Cost

The second merit index was cost (3). 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 raw material cost for the alloying element, c,.

Cost = åi XiCi (2)

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

Heat Treatment Window

A third merit index is the solutioning window (4). By conducting thermodynamic modelling (CALPHAD) calculations across a range of temperatures the solutioning window for each alloy can be calculated. This value - measured in degrees Celsius - can be used to determine if a given alloy is amenable to conventional manufacturing processes used for the production of single crystal turbine blades. Typically the solutioning window should be greater than 50°C to allow for a solution heat treatment. The solution heat treatment is conducted in the single phase region, at this point the alloy will reside solely within the g phase field. This solution heat treatment is necessary to homogenise the composition of the as cast alloy which may be highly segregated. In order determine the solution heat treatment window the phase equilibrium - or more specifically phase transformations - must be determined over a temperature range. The temperature at which completed dissolution of the g’ phase (known as the g’ solvus temperature) occurs must be known, as must the solidus temperature. The difference between the solidus temperature and the g’ solvus temperature will give the solutioning window. So the solutioning window index calculates as the difference between the solidus temperature and the g’ solvus temperature.

TCP Phases:

The presence of TCP phases (m, s & P) is detrimental for long term mechanical properties in Ni-based superalloys. Their appeared is cause by long high temperature exposition and they are greatly influence by the alloy chemistry. It has been found that the propensity of forming TCP phases is directly linked with the d-orbital energy levels of the g composition in the alloy (Md_Y). From experimental observation it has been observed that alloys with Md_Y < 0.93 do not form TCP phases. In this invention, the Md_Y values (5) for each alloy composition has been calculated using thermodynamical databases. 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

W_d = å_i x_iMd_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.

Oxidation:

Oxidation of superalloys depends strongly on the oxides layers form on their surface. Chromia and Alumina (AI2O3) layer has been proved to be beneficial for the oxidation resistance of the alloy. Chromia is formed when sufficient amount of chromium is achieved in the alloy. However, for SX this amount is not sufficient due to the creep requirement. On the other side, alumina formation, which is the main barrier in most commercial SX alloys, is promoted when low effective valences of AI2O3 (Val^eff) and low formation energies of this compound (AG^f) are combined together. Following this, works in the literature have shown that there is a direct correlation between the oxidation constant k_t and the propensity of the alloy to oxidise, k_t defined as:

k_t=10 (Val^eff · AG^f-k₀) where ko (481.6 kJ/moP is a constant threshold to form protective alumina. We have used this oxidation constant (6) as the sixth merit index to rank alloys for oxidation resistance. The parameters Val^eff and AG^f can be obtained from thermodynamics databases dependent on the composition. The index (k_t) is more positive when the alloy is less oxidation resistant. From experimental results in the literature, k_t is typically around -0.10 for the 3^rd generation superalloys and worse for 4^th generation. In this work, we use k_t = -0.10 as our aimed target.

Mechanical Indexes

Yield Strength (0-600°C)

The seventh merit index (7) is the yield strength. This yield strength is derived from a physical based model combining the amount of g’ fraction (f/) and the strength of these precipitates.

Where T is the Taylor factor (for SX <001> is 1/0.41), b is the burger vector and g_ARB is the antiphase boundary (APB) energy. The g’ fraction is obtained as stated before from thermodynamic databases. The fault energies in the g' phase - for example, including g_ARB- 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 g' 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 = 195— 1.7x_Cr— 1.7X_MO + 4.6x_w + 27.1c_Ta + 21 Ax_Nb + 15c_tί (4) where, xcr, XM_O, XW, Cta, xm and xi, represent the concentrations, in atomic percent, of Cr, Mo, W, Ta, Nb and Ti in the g' phase, respectively. The composition of the g' phase is determined from phase equilibrium calculations.

Mid-temyeratures strength ( 600-800°C )

The eigth merit index (8) is the mid-temperature Larson Miller Parameter (LMP_MUI) and define the resistance of the alloy to mid-temperature deformation (600-800°C). The definition of the LMP is:

(Where T is the temperature condition and t_10o is the time to reach 1% of plastic deformation. This LMP depends strongly on the stress condition (s) and the chemical composition of the alloy. The dependence of the LMP_Mid on the chemical composition of the alloy is obtained by a combined used of physical observation in the literature and machine learning. The response of a database of 150 experimental data points of the mid-temperature creep response of 39 alloys is used to train the model. The linear equation obtained from the machine learning process is:

LMP_Mid = 11.4— 0.1 x_Cr— 0.2 x_Co + 0.3 x_Ta + 0.3 x_w— 0.08 x_Re + 0.66 x_Ru + 20.76 /y - 1.16 log (s - 500) (6)

These linear coefficients are aligned with recent discoveries of the fault shearing mechanism active in this range. These studies found that Co and Cr accelerate the growth of these faults while Ta, W and g' to limits their extension rate thus improving the creep strength of the alloy. The accuracy for the computed LMP_Mid is presented in Fig. 13. To liberate the optimisation process from the stress variable a typical stress of 650 MPa for this temperature range is fixed for all the alloy compositions.

High-temyeratures strength ( 600-800°C )

The last merit index is the creep-merit index (9). The overarching observation is that creep deformation of a single crystal superalloy above 800°C occurs by dislocation climb of the g' which is highly dependent on the alloy chemistry. Because of this change of mechanistic above 800°C a new chemical composition dependence is being calculated for LMP c_reep. In this case, a database of 1314 experimental datapoints and 120 alloys is used to train the linear model using machine learning. The chemical linear relation found for LMP c_reep is:

LMP_Creep = 22.81— 0.08 x_Co + 0.08 x_Ta + 0.13 x_w + 0.19 x_Re + 0.26 x_Ru—

0.015 s (7)

Again, these linear coefficients are aligned to what is expected from a physical point of view. Many studies reveal the strong effect of Re, Ru and W on the creep resistance of Ni- based superalloys. The accuracy for the computed LMP_Mid is presented in Fig. 14. To liberate the optimisation process from the stress variable a typical stress of 250 MPa for this temperature range is fixed for all the alloy compositions.

The ABD method described above was used to isolate the inventive alloy composition. The design intent for this alloy was to isolate the composition of a fourth-generation single crystal (SX) nickel-based superalloy that exhibits a superior creep resistance which is comparable or better than equivalent grades of alloy such as TMS-138A but keeping the strength, stability and oxidation and corrosion resistance of a 2^nd generation SX. The density, cost, processing of the alloy have also been considered in the design of the new alloy with similar values to the ones obtained for 4^th generation SX.

The material properties - determined using the ABD method - for the commercially used fourth generation single crystal turbine blade alloys are is listed in Table 3. The design constrains of the new alloy were established by the predicted values of baselines alloys (HT creep, cost, density, HT window, MT strength) and 2^nd generation alloys (yield strength, oxidation and corrosion resistance, stability) as stated in the final row of Table 3. As can be seen the wide variety of alloys falling in the scope of the invention meet all of the criteria set in the final row of table 3, whereas each of the prior art alloys does not meet at least at least two criteria. In particular the alloys of the present invention achieve good creep resistance at high temperatures and mid-temperatures in combination with a high yield stress and high yield stress in combination with a solutioning window exceeding 50 °C. The calculated material properties for a set of example alloys in accordance with the present invention are also given. The composition for the example set of alloys are stated in Table 5. Table 3: Calculated phase fractions and merit indices made with the“Alloys-by-Design” software. Results for fourth generation single crystal turbine blades listed in Table 1 and the proposed compositions in the current invention listed in Table 5.

Microstructural Optimisation:

Optimisation of the alloy’s microstructure - primarily comprised of an austenitic face centre cubic (FCC) gamma phase (g) and the ordered Lb precipitate phase (g') - was required to maximise creep resistance. A volume fraction of the g' phase between 60-70% is generally regarded as optimum as this microstructure is known to provide the maximum level of creep resistance in single crystal blade alloys. A volume fraction g' of between 60 and 70% was the target for the present alloy but the inventive alloy may deviate from this target. Aluminium and tantalum are well known to be the primary g’ formers. Hence, the levels of these elements were controlled to produce the desired g' volume fraction. Figure 3 shows the effect the elements which are added to form the g' phase - predominantly aluminium and tantalum - have on the fraction of g' phase in the alloy at the operation temperature, 900°C in this instance. For the design of this alloy compositions which result in a volume fraction of g' between 60-70% were considered. Hence between 5.3 and 6.8 weight percent (wt.%) of aluminium was required based on a tantalum required content of 6.0 - 9.7 wt% (described with reference to figures 8 and 10).

The change in g' volume fraction was related to the change in aluminium and tantalum content according to the formula

/(y') = w_{Ta +} 5.0 w where, f(y) is a numerical value which ranges between 34 and 39 for an alloy with the desired g' fraction, between 0.6 and 0.7 in this case, and Wm and WAI are the weight percent of tantalum and aluminium in the alloy, respectively.

Optimisation of aluminium and tantalum levels was also required to increase the yield strength of the alloy. Figure 9 shows the influence of aluminium and tantalum on the yield strength of the alloy. The typical yield strength of fourth generation alloys is lower than for third generation alloys which can lead to low cycle fatigue problems. The typical yield strength of a third generation alloy is aimed at in the present invention combined with a fourth- generation creep resistance. Compositions where the yield strength achieves the strength of third generation single crystal alloys (-945 MPa) are indicated in the graph. Modelling calculation showed that tantalum levels in the alloy greater than 6.0 wt.% is needed to produce an alloy with an acceptable yield strength at the lowest level of aluminium allowable. Desirably the alloy has at least 6.5 wt% tantalum so that even at the highest aluminium levels allowable the yield stress is comparable to the 3^rd generation alloys. The yield strength can be related by the formula

f(YS = W_Ta + 0A6W_AI where, f(YS) is higher than 3.34 for a YS of 945 MPa.

The maximum tantalum content will be explained below with reference to Figure 7 and results in a tantalum range of 6.0 - 9.7 wt.%. A preferred range of 6.5 to 8.0 wt.% results from the preferred combination of yield strength and solutioning window (dealt with below). That is, the preferred minimum levels of tantalum ensure a high enough yield strength for any given amount of aluminium and a level of945 MPa in the range of aluminium for the alloy and the max. amount of Ta=8.0% assures an admissible HT window for the preferred range of Cr (Fig. 8 - justified later). From Figure 3 it is seen that for the new preferred range of tantalum, concentrations of aluminium between 5.5% and 6.6% produce the desired volume fraction of g'. Therefore, it is preferable to have the ratio of aluminium to tantalum, in weight percent, ranging between 0.54 (Al=5.3 wt.%, Ta=9.3 wt.%) and 1.13 (Al=6.8 wt.%, Ta=6.0 wt.%), or more preferably ranging between 0.68 (Al=5.5 wt.%, Ta=8 wt.%) and 1.02 (Al=6.6 wt.%, Ta=6.5 wt.%).

Niobium, titanium and vanadium behave in a similar way to that of tantalum i.e. they are gamma prime forming elements which increase anti-phase boundary energy. These elements can optionally be added to the alloy. The benefits of this may include lower cost and density in comparison to tantalum. However, additions of these elements must be limited as they can have a negative impact on the environmental resistance of the alloy. Therefore, those elements can each be present in an amount of up to 0.5 wt.%. Preferably those elements are substituted for tantalum meaning that the sum of the elements consisting of niobium, titanium, vanadium and tantalum is preferably limited to 6.0-9.7 wt.%, more preferably 6.2-8.0 wt.% which is the preferred range for tantalum. Independently, in an embodiment, the sum of the elements consisting of niobium, titanium and vanadium is preferably limited to below 1.0 wt.% and preferably below 0.5 wt.% so as to avoid reduction in environmental resistance of the alloy.

The elements platinum and palladium behave in a similar way to that of tantalum, titanium and niobium i.e. they are g' forming elements which increase anti-phase boundary energy. These elements can optionally be added to the alloy for example in substitution for the elements tantalum, titanium, vanadium and niobium. The benefits of this may include an improvement in resistance to high temperature corrosion. However, additions of these elements can be limited due to the high cost of these elemental additions. Therefore, those elements can each be present in an amount of up to 1.0 wt% or less and most preferably 0.5 wt% or less as this range provides the best balance of cost and improvement to corrosion resistance.

The element iridium behaves in a similar way to that of tungsten i.e. it is a gamma forming element which improves the creep merit index. Iridium can optionally be added to the alloy. Additions of iridium will significantly increase the creep response of the alloy in comparison to tungsten (as it has much slower diffusivity), however this is achieved with substantial increases in cost due to the high cost of iridium. Preferably the addition of iridium is limited to 1.0 wt% or less and even more preferably to 0.5 wt% or less.

The balance of aluminium and tantalum can be adjusted such that there is a balance between desired target volume fraction of g' as well as a sufficiently high yield strength. However, consideration must also be given to the processing of the alloy. One such consideration is the solutioning window; there should exist a sufficient temperature range window, below the melting temperature of the alloy, across which only the g phase is stable. As the solutioning window depends upon the dissolution of the g' phase it is strongly influenced by alloy chemistry, specially by tantalum and chromium content. This solutioning heat treatment is used to remove any residual microsegregation and eutectic mixtures rich in g' which might occur during the casting processes used to produce the single crystal alloy. It is preferred that the solutioning window is greater than 50°C to allow for conventional processing methods. Figure 7 shows the solutioning window magnitude (in °C) for varying wt% Cr and Ta with a volume fraction g' of 60-70%. The minimum chromium content for the present invention is greater than or equal to 3.0 wt.% and preferably greater than or equal to 4.0 wt.% in order to attain oxidation resistance which is improved in comparison to current fourth generation single crystal alloys which have Cr contents ranging between 2.0-3.2 wt.%. That is, a higher weight percent of chromium is provided than in the current fourth generation alloys on the basis that this will improve oxidation resistance compared to those alloys. The chromium content is limited to 6.5 wt.% to reduce the propensity for the alloy to form the deleterious TCP phases which will be explain later (Figure 6). Preferably the chromium content in the alloy is limited to 5.0 wt.% as this produces an alloy with the best balance between oxidation resistance and microstructural stability. With the maximum allowable amount of Cr of 6.5wt%, from Figure 8 it can be seen that limiting the tantalum content to 9.7 wt.% ensures that the alloy has a suitable solutioning window. Preferably the tantalum content is limited to 8.0 wt.% as this produces an alloy with a solutioning window greater than 50°C even at lower levels of Cr. Limiting tantalum even further to say 7.0 wt.% or less may be beneficial for further increasing the solutioning window.

The change in solutioning window was related to the change in chromium and tantalum content according to the formula f sol. ) = W cr - W _Ta where, f(T_soi) is a numerical value which is greater than or equal to 3.5 to produce an alloy with a solutioning window greater than or equal to 50°C. f(T_soi) is preferably greater than or equal to 3.0 to produce and alloy with a solutioning window greater than 60°C.

For the alloys which satisfied the previously described requirements (volume fraction of g’ between 60-70%, yield strength greater than 945 MPa, solutioning window greater than 50°C) the levels of refractory elements were determined for the required creep resistance and mid-temperature strength. For fourth generation single crystal turbine blades additions of the elements ruthenium, rhenium and tungsten (in order of importance) are made to impart substantial creep performance, this is described later with reference to Figure 11. However, the elements rhenium and ruthenium strongly affect cost, Figure 5. The elements tungsten and rhenium significantly increase alloy density, Figure 4. Moreover, elements such as rhenium, tungsten and chromium (chromium is added for oxidation resistance as stated before) must be suitably balanced such that a balance between creep resistance and oxidation is achieved without resulting in a microstructurally unstable alloy which is prone to the formation of deleterious TCP phases, Figure 6.

As Cr content has been fixed between 3.0-6.5 wt. %, the Re + W content must be carefully tailored to provide the desired creep performance without making that alloy unstable. Thus, a complex balance between trade-offs in cost, density, creep resistance, oxidation resistance and microstructural stability must be managed, the process for optimising these trade-offs is described below with reference to Figures 4, 5, 6,1 1.

The current raw material cost for the elements ruthenium and rhenium is substantial. Therefore, to optimise the design of the alloy levels of ruthenium and rhenium are selected which best manage the trade-off between the cost and creep resistance in the present invention. In Figure 5 a contour plot shows the effect which levels of ruthenium and rhenium have on alloy cost for an alloy of 60-70% g' volume fraction at 900°C. It is seen that ruthenium has the strongest influence on alloy cost. Thus, the ruthenium content in the alloy is limited to 4.1 wt.% ensuring that the cost of the present invention is about 340 $/lb at the minimum allowed amount of rhenium of 3.5wt%, explained elsewhere. A preferred maximum allowable amount of ruthenium is 4.0wt% as this keeps to cost to about equivalent to or less than current grades of fourth generation alloy (325$/lb). It is preferred that the ruthenium content is limited to 3.5 wt% to ensure an optimal balance between cost and creep resistance as indicated by the small design region in Fig l id.

In order to limit the cost of the alloy, as an approximation additions of ruthenium and rhenium preferably adhere to the following Equation,

/(Cost ) = W_Ru + 0.22 W_Re where, /(Cost) is a numerical value which is less than or equal to 4.8 to produce an alloy with a cost of 325$/lb or less and W_RU and W_RC is the weight percent of ruthenium and rhenium in the alloy respectively and so this is a preferable feature. Preferably the numerical value for /(Cost) is less than or equal to 3.7 as this produces an alloy with a lower cost of 300$/lb or less.

The additions of the elements tungsten, rhenium and ruthenium are optimised in order to design an alloy which is highly resistant to creep deformation. The high temperature creep resistance was determined by using the LMP_creep model developed using machine learning tools using a first order function of the chemical compositions. The model is trained with 1047 experimental points mined from the literature. The model is then used to predict the LMP_creep. The correlation between experimental values and predicted values by the model is presented in Fig. 14. It is desirable to maximise the LMPcreep as this is associated with an improved high temperature (HT) creep resistance. The influence which tungsten, rhenium and ruthenium have on the HT creep resistance is presented in Figure 1 1. The calculations to produce the graphs of Figures 4 and 11 are with the limitation that the g' volume fraction at 900°C is between 60 and 70% enforced. It is seen that increasing the levels of ruthenium, rhenium and tungsten improve creep resistance, by this order of effectiveness. However, the quantities of tungsten and rhenium required mean that they have a strong influence on alloy density, Figure 4. Therefore, the trade-off between creep resistance and alloy density must be balanced.

In order to limit the density of the alloy additions of tungsten and rhenium preferably adhere to the following Equation,

f (Density) = 1.09 W_Re + W_w where, /(Density) is a numerical value which is less than or equal to 13.6 to produce an alloy with a density of 9.0 g/cm³ or less and Ww is the weight percent of tungsten in the alloy. Preferably the numerical value for /(Density) is less than or equal to 11.6 as this produces an alloy with a density of 8.9 g/cm³ or less.

Current fourth generation single crystal alloys have a LMP_creep of 25.6 or greater (see Table 3), or even better 25.7 ( TMS-138A). This level of high temperature creep resistance is desirably attained in combination with a low density of less than 9.0 g/cm³ or preferably 8.9 g/cm³. In Figure 11 the density contours from Figure 4 (dashed lines) and the cost contours from Figure 5 (dot lines) are superimposed on the effect which rhenium, tungsten and ruthenium have on LMPcreep.

In order to attain a LMPcreep of 25.6 without having large qualities of rhenium, the alloy contains at least 2.0 wt.% of ruthenium. Preferably the ruthenium content is greater as this clearly increases LMPcreep, so that preferably ruthenium is present in an amount of 2.5wt% or more, more preferably 2.8wt% or more and even more preferably 3.0 wt.% or greater as this produces even higher creep resistance, i.e. LMPcreep > 25.7. Ruthenium is limited to 4.1 wt.% as this gives the preferred balance between cost and creep resistance. Preferably ruthenium is limited to 4.0wt% as this gives an even better balance between cost and creep resistance.

The tungsten content is limited to 7.5 wt.% or less, so that the alloy density can preferably be decreased to 9.12 g/cm³ or less (equivalent to PW1497 and MX4) at maximum rhenium content of 8.5% (figure 4). Preferably the tungsten content is limited to 6.0 wt% or less or even 5.2wt% or less or even 5.0 wt.% or less as this produces an alloy with an even lower density (dashes lines in Figures 4 and 11c). Lower levels of tungsten also ensure microstructural stability (Figure 7).

From Figure 11 a minimum content of rhenium of 3.5 wt.% or more is shown to produce a LMPcreep of 25.6 at the maximum allowable amounts of ruthenium of 4. lwt% (or 4.0wt%) and tungsten of 7.5wt%. Preferably the rhenium content is greater than 5.0 wt.% as this produces an alloy with a better balance between density (Figure 4) and creep resistance (Figure 11c). Even more preferable is an alloy containing at least 6.5 wt.% of rhenium as this composition produces an alloy with an even better balance of creep resistance and density. In such an alloy cost can also be reduced as lower levels of ruthenium may be required than with lower rhenium levels (Figure 1 lc-d). Molybdenum behaves in a similar way to tungsten i.e. this slow diffusing element can improve creep resistance. Therefore, although molybdenum additions are optional, it is preferred that molybdenum is present in an amount of at least 0.1 wt%. However, additions of molybdenum must be controlled as it strongly increases the alloys propensity to form deleterious TCP phases. Therefore, molybdenum is limited to 0.5 wt.% or less.

From Figures 1 la-d and a knowledge that molybdenum can substitute tungsten, it can be determined that a good level of creep resistance is achieved when additions of tungsten, rhenium, molybdenum adhere to the following Equation,

/(Creep) = 0.29 W_Ru + 0.1W_Re + 0.026(W_w + W_Mo ) where, W_MO is the weight percent of molybdenum in the alloy and /(Creep) is a numerical value. If /(Creep) is greater than or equal to 1.81, this produces an alloy with a LMP_creep.as calculated of 25.6. Preferably the numerical value for /(Creep) is greater than 1.86 as this produces an alloy with increased creep resistance similar to TMS-138A.

In order to remain resistant to creep over a significant time period the addition of slow diffusing elements rhenium, tungsten and ruthenium is required. Additions of chromium are also required to promote resistance to oxidation/corrosion damage. However, the addition of high levels of tungsten, rhenium and chromium were found to increase the propensity to form unwanted TCP phases, primarily s, P and m phases. Figure 6 shows the effect of chromium, tungsten and rhenium additions on effective valence MD·, (higher valence, more prone to form TCP phases). Preferably the additions of these elements are controlled to ensure a value MD_y<0 93 which has been found to be the barrier for TCP phases formation. All alloys selected in this invention are just below this limit in Table 3. With this in mind and accounting for a 3.0-6.5 range in chromium, in the present invention the rhenium content in the alloy is limited to 8.5 wt.% or less (to ensure acceptable microstructural stability, Figure 6d-e particularly for low levels of tungsten) and more preferably 8.0 wt.% or less as rhenium at a level of between 6.0 wt.% and 8.0 wt.% provides a good balance between density, creep resistance and microstructural stability. Good microstructural stability is assured if the following equation is satisfied:

W_w+1.2714^-0.83 W_Re<\ 8.2 Based upon the rhenium levels allowable (W_Re< 8.5%) at the minimum allowable level of ruthenium of 2.0 wt% for a balance between microstructural stability, density and creep resistance the minimum tungsten level required for the present invention is 1.5 wt.% or more, as this provides a balance between creep resistance (Figure 11) and cost (in terms of reducing necessary levels of rhenium and ruthenium) without reducing microstructural stability too much (Figure 6). In order to achieve high creep resistance, i.e. LMP_creep of 25.7 (Figure 11), a preferred minimum level of tungsten is 2.0 wt.%.

The last element to define is cobalt. In Ni-based superalloys, cobalt has been observed to have a strong beneficial effect on the oxidation and corrosion providing a defence against sulphidication. However, in the recent years, cobalt has been found to promote the plastic deformation at mid-temperature, thus reducing the strength of the alloy between 600-800°C. Cobalt cost is also a highly fluctuating in the last few years so limitation on the amount of this element would be beneficial from an economical point of view. There is a clear trade-off for cobalt between oxidation, mid-temperature strength and cost. Figure 8 shows the effect of Co and W on the oxidation index. Additions of cobalt improve the oxidation behaviour of the alloy and allows for higher amounts of W. A critical threshold of the oxidation index of -0.1 can be imposed to have a similar good oxidation resistance than 2^nd generation SX alloys (better than 4^th generation). From this figure, for the minimum amount of tungsten (Ww>1.5%) the minimum value of Cobalt for a good oxidation resistance is set to 4.0%. Preferably, alloys with cobalt higher than 5.0% present a better oxidation resistance. Further increasing cobalt levels increases oxidation levels further and the examples show that excellent properties can be maintained. Thus preferred minimum amounts of cobalt are 7.5wt% or more, more preferably 7.8wt% or more or most preferably 8.0wt% or more.

0.63Wco - Ww > -3.6;

preferably 0.63Wco - Ww > 0.83

On the other hand, the cobalt amount is limited by the propensity of the alloy to form low energy continuous fault thus ruining the strength of the alloy at mid-temperature. The effect of cobalt on the strength of the alloy at mid temperatures is presented in Figure 10 for alloy with 60-70% g’ fraction and typical values of Cr (4.5-6 wt. %) and Ta (6-8.5 wt. %). From the model used, a good mid-temperature strength can be achieved following this equation:

/ (Mid - Temp ) = 1.10 (W_Co + W_Cr) - (W_Ta + W_w) Where the function /is preferably lower than 2.15 to assure a good strength at mid temperature with a LMP_MUI of 21.5 or greater. For the maximum levels of tungsten allowed in this invention (7.5 wt. %) and the ranges of Cr and Ta defined, the maximum cobalt level to keep a good strength at mid temperatures is 12.0% and preferably 10.5% or more preferably 10.3% or even more preferably 10.0% for an improved strength.

It is beneficial that when the alloy is produced, it is substantially free from incidental impurities. These impurities may include the elements carbon (C), boron (B), sulphur (S), zirconium (Zr) and manganese (Mn). If concentrations of carbon remain at 100 PPM or below (in terms of mass) the formation of unwanted carbide phases will not occur. Boron content is desirably limited to 50 PPM or less (in terms of mass) so that formation of unwanted boride phases will not occur. Carbide and boride phases tie up elements such as tungsten or tantalum which are added to provide strength to the g and g' phases. Hence, mechanical properties including creep resistance are reduced if carbon and boron are present in greater amounts. The elements Sulphur (S) and Zirconium (Zr) preferably remain below 30 and 500 PPM (in terms of mass), respectively. Manganese (Mn) is an incidental impurity which is preferably limited to 0.05wt% (500PPM in terms of mass). 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. This segregation may lead to increased spallation of protective oxide scales. The levels of zirconium and manganese must be controlled as these may create casting defects during the casting process, for example freckling. If the concentrations of these incidental impurities exceed the specified levels, issues surround product yield and deterioration of the material properties of the alloy is expected. Copper is an incidental impurity which is preferably limited to 0.5 wt%.

Additions of hafnium (Hf) of up to 0.5wt.%, or more preferably up to 0.2wt.% are beneficial for tying up incidental impurities in the alloy, in particular carbon. Hafnium is a strong carbide former, so addition of this element is beneficial as it will tie up any residual carbon impurities which may be in the alloy. It can also provide additional grain boundary strengthening, which is beneficial when low angle boundaries are introduced in the alloy.

Iron behaves in a similar way to nickel and can be added as a low-cost alternative to nickel. Moreover, tolerance to iron additions improves the ability of the alloy to be manufactured from recycled materials. Therefore, it is preferred that iron is present in an amount of at least 0.1 wt%. However, additions of iron up to 4.0 wt% can be made in order to substantially reduce the cost. Preferably the additions of iron are 2.0 wt% or less in order to reduce the propensity to form the unwanted Laves phase which degrades the mechanical properties of the alloy. Most preferably iron additions are limited to 1.0 wt% as this produces an alloy which has good ability to be recycled with no loss in material performance.

Additions of the so called‘reactive-elements’, Silicon (Si), 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 AI2O3. 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. In particular, it has been shown that additions of silicon to nickel based superalloys at levels up to 0.1 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. Magnesium (Mg) likewise may act to‘mop up’ tramp elements, and can have beneficial effects on mechanical properties, so may be added up to 0.1%.

Based upon the description of the invention presented in this section, broad and preferred ranges for each elemental addition were defined, these ranges are listed in Table 4. The alloys satisfying all the requirement imposed previously where ranked in terms of their mechanical properties in a multitarget optimisation exercise. This is presented in Figure 13 as a triangular graph with each vertex a mechanical property (yield strength, mid-temperature strength and creep strength). The perfect alloy would fill the whole triangle while satisfying all the other requirements. The example alloys for this invention were then selected by the ranking the final set of alloys from the degree of fulfilment of the vertex of the triangle. Some examples of 2^nd ,3^rd and 4^th generation superalloys are presented along with examples alloys in Figure. 13. The compositions of all the example alloys selected are presented in Table 5. Table 4: Compositional range in wt.% for the newly design alloy.

Broad Preferred Most Preferred

Min Max Min Max Min Max

Cr 3.0 6.5 3.0 6.5 4.0 5.0

Co 4.0 12.0 7.8 12.0 7.8 10.5

W 1.5 7.5 1.5 5.2 2.0 5.2

A1 5.3 6.8 5.3 6.8 5.5 6.6

Ta 6.0 9.7 6.0 9.7 6.5 8.0

Mo 0.0 0.5 0.0 0.5 0.1 0.5

Re 3.5 8.5 3.5 8.5 5.0 8.0

2.5 or 2.8 4.1

Ru 2.0 4.1 or 2.5 or 2.8 3.5

4.0

Hf 0.0 0.5 0.0 0.5 0.0 0.2

Nb 0.0 0.5 0.0 0.5 0.0 0.5

Ti 0.0 0.5 0.0 0.5 0.0 0.5

V 0.0 0.5 0.0 0.5 0.0 0.5 Si 0.0 0.1 0.0 0.1 0.0 0.1 Ce 0.0 0.1 0.0 0.1 0.0 0.1

Y 0.0 0.1 0.0 0.1 0.0 0.1 La 0.0 0.1 0.0 0.1 0.0 0.1

Table 5: Compositions of the example alloys proposed in this invention n wt. %.

Claims

1. A nickel-based alloy composition consisting, in weight percent, of: 3.0 to 6.5% chromium, 4.0 to 12.0% cobalt, 1.5 to 7.5% tungsten, 0.0 to 0.5% molybdenum, 3.5 to 8.5% rhenium, 2.0 to 4.1% ruthenium, 5.3 to 6.8% aluminium, 6.0 to 9.7% tantalum, 0.0 to 0.5% hafnium, 0.0 to 0.5% niobium, 0.0 to 0.5% titanium, 0.0 to 0.5% vanadium, 0.0 to 1.0 platinum, 0.0 to 1.0 palladium, 0.0 to 1.0 iridium, 0.0 to 0.1% silicon, 0.0 to 0.1% yttrium, 0.0 to 0.1% lanthanum, 0.0 to 0.1% cerium, 0.0 to 0.1% magnesium, 0.0 to 0.003% sulphur, 0.0 to 0.05% manganese, 0.0 to 0.05% zirconium, 0.0 to 0.005% boron, 0.0 to 0.01% carbon, 0.0 to 0.5% copper, 0.0 to 4.0% iron, the balance being nickel and incidental impurities.

2. The nickel -based alloy composition according to claim 1, consisting, in weight percent, of 5.0% or less chromium.

3. The nickel -based alloy composition according to claim 1 or 2, consisting, in weight percent, of 5.0% or more cobalt, preferably 7.5% or more cobalt, more preferably 7.8% or more cobalt, even more preferably 8.0% or more cobalt.

4. The nickel-based alloy composition according to claim 1, 2 or 3, consisting, in weight percent, of 10.5% or less cobalt, preferably 10.3% or less cobalt, more preferably 10.0% or less cobalt.

5. The nickel -based alloy composition according to claim 1, 2, 3 or 4, consisting, in weight percent, of 2.0% or more tungsten.

6. The nickel -based alloy composition according to any of claims 1-5, consisting, in weight percent, of 6.0 or less tungsten, preferably 5.2 or less tungsten, more preferably 5.0 or less tungsten.

7. The nickel -based alloy composition according to any of claims 1-6, consisting, in weight percent of, 6.6% or less aluminium.

8. The nickel-based alloy composition according to any of claims 1-5, consisting, in weight percent, of 8.0% or less tantalum, preferably of 7.0% or less tantalum .

9. The nickel -based alloy composition according to any of claims 1-5, consisting, in weight percent, of 6.5% or more tantalum.

10. The nickel -based alloy according to any of claims 1-9, consisting, in weight percent, of at least 0.1% molybdenum.

11. The nickel-based alloy according to any of claims 1-10, consisting, in weight percent, of 5.0% or more rhenium, preferably 6.5% or more rhenium.

12. The nickel-based alloy according to any of claims 1-11, consisting, in weight percent, of 8.0% or less rhenium.

13. The nickel-based alloy according to any of claims 1-12, consisting, in weight percent, of 2.5% or more ruthenium, preferably 2.8% or more ruthenium, more preferably 3.0% or more ruthenium.

14. The nickel-based alloy according to any of claims 1-13, consisting, in weight percent, of 4.0% or less ruthenium, preferably 3.5% or less ruthenium.

15. The nickel-based alloy composition according to any of claims 1-14, wherein the following equation is satisfied in which W_T3 and WAI are the weight percent of tantalum and aluminium in the alloy respectively

34 < W_Ta + 5.0 WAI £ 39.

16. The nickel-based alloy composition according to any of claims 1-15, wherein the following equation is satisfied in which W_RU and Wit_e are the weight percent of ruthenium and rhenium in the alloy respectively

4.8 > WR_U + 0.22 W_Re,

preferably 3.7 > W_Ru + 0.22 W_Re.

17. The nickel-based alloy composition according to any of claims 1-16, wherein the following equation is satisfied in which W_Re and Ww are the weight percent of rhenium and tungsten in the alloy respectively

13.6 > 1.09 W_Re + W_w preferably 11.6 > 1.09 Wit_e + Ww.

18. The nickel -based alloy composition according to any of claims 1-17, wherein the following equation is satisfied in which Wit_e, W_MO and Ww are the weight percent of rhenium, molybdenum and tungsten in the alloy respectively

1.81 < 0.29 W_Ru + 0.1 W_Re + 0.026(W_w + W_Mo), preferably 1.86 <0.29W_Ru + 0.1W_Re + 0.026(W_w + W_Mo).

19. The nickel-based alloy composition according to any of claims 1-18, wherein the following equation is satisfied in which WAI and Wx_a are the weight percent of aluminium and tantalum in the alloy respectively

0.46 WAI + W_Ta > 3.34

20. The nickel-based alloy composition according to any of claims 1-19, wherein the following equation is satisfied in which Ww_, Wc_o, Wc_r and Wx_a are the weight percent of tungsten, cobalt, chromium and tantalum in the alloy respectively

1.10(Wco + Wcr) - (Wxa + Ww) < 2.15

21. The nickel -based alloy composition according to any of claims 1-20, wherein the following equation is satisfied in which Wc_r and Wx_a are the weight percent of chromium and tantalum in the alloy respectively

3.5 < W Cr - W _Ta

preferably 3.0 < Wcr - Wx_a

22. The nickel -based alloy composition according to any of claims 1-21, wherein the following equation is satisfied in which Wc_o and Ww are the weight percent of cobalt and tungstem in the alloy respectively

0.63 Wco - Ww > -3.6;

preferably 0.63Wco - Ww > 0.83

23. The nickel -based alloy composition according to any of claims 1-22, wherein the following equation is satisfied in which Wc_r, Wit_e and Ww are the weight percent of chromium, rhenium and tungstem in the alloy respectively

W_w+l .27W_Cr-0.S3W_Re<lS2

24. The nickel-based alloy composition according to any of claims 1-23, wherein the sum of the elements niobium, titanium and vanadium, in weight percent, is less than 1%, preferably less than 0.5 wt%.

25. The nickel -based alloy composition according to any of claims 1-24, consisting, in weight percent, of between 0.0 and 0.2% hafnium.

26. The nickel -based alloy composition according to any of claims 1-25, consisting, in weight percent of, 5.5% or more aluminium.

27. The nickel -based alloy composition according to any of claims 1-26, having between 60% and 70% volume fraction g'.

28. The nickel -based alloy according to any of claims 1-27, wherein the sum of the elements niobium, titanium, vanadium and tantalum, in weight percent, is between 6.0 and 9.7%, preferably between 6.5 - 8.0%.

29. The nickel -based alloy according to any of claims 1-28, consisting of, in weight percent, 4.0 % or more chromium.

30. A single crystal article formed of the nickel-based alloy composition of any of claims 1-29.

31. A turbine blade for a gas turbine engine formed of an alloy according to any of claims 1-29.

32. A gas turbine engine comprising the turbine blade of claim 31.