Introduction

Gas turbines are widely used in power generation sys tems. Developments to improve their efficiencies have  led to increased operating temperatures of regions of some components, such as the under platform areas of turbine blades. The high stress state of the root pocket due to the high rotational speeds, combined with cooling air derived deposits and temperatures approaching the conditions associated with type II corrosion, can lead to cracking [1].

CMSX-4 (Table 1) is a single crystal Ni-based superalloy commonly used for 1st stage gas tur bine blades as a result of its good high temperature creep-strength properties combined with production affordability [2]. However due to its composition (lower Cr content than other commonly used 1st stage turbine blade materials), CMSX-4 is susceptible to type II hot corrosion. This can result in damage that has the morphology of either pitting or broad fronted attack. Sumner et al. [3] have reported investigations of type II hot corrosion of CMSX-4, using statistical analysis of large data sets to generate models for spe cific conditions. They observed broad fronted attack and more rapid depletion of Cr in CMSX-4, when compared with IN738LC.

      Reaserch conducted on hot corrosion mechanisms in 1970s–80s is summarised by Luthra & LeBlanc [4]. They concluded that hot corrosion could occur through a combination of three mechanisms: sulphidation-oxidation, formation of volatile compounds beneath the protective oxide layer or scale fluxing. Fluxing models have since gained the widest acceptance for deposit induced hot corrosion [5,6]. 

        The process of type II hot corrosion of Ni-based superalloys requires the formation of a liquid eutectic film [5,6]. Type II hot corrosion occurs in the temper ature range of 650–800 °C through the formation of minimum melting point mixtures of Na2 SO4 , NiSO4 and CoSO4 [4,5,8]. NiSO4 and CoSO4 compounds form as a result of the reaction of SO3 with nickel and cobalt from the superalloy. A widely accepted mechanism for hot corrosion was proposed by Goebel & Pettit [9]. Their mechanism outlines two stages, firstly the incu bation stage, where a liquid eutectic of Na2 SO4 , NiSO4 and/or CoSO4 forms on the component surface as a result of deposition coupled with a reaction between sulphur oxides and nickel and/or cobalt from the super alloy. The second stage is the propagation stage, where the fluxing of the surface oxide by a liquid deposit on the surface allows inward access, and outward Co/Ni  transport. This form of attack often results in pitting damage with an outer NiO/CoO layer being formed, although sometimes a form of broad fronted attack develops [5,6].

For type II hot corrosion many researchers have noted the importance of a constant SOx supply for sustained corrosion to occur [3,7,9,10]; this variation of the damage mechanism is known as gas induced acidic fluxing [8,11]. Without both gaseous SOx and a regular sulphate deposition flux, the corrosion reac tion would cease to occur when all the reactants have been consumed.

Type II hot corrosion combined with static stress in Ni-based superalloys has not been extensively studied. However, stress corrosion cracking (SCC) is a well-documented failure mechanism especially in aqueous systems [12,13].

Studies have been conducted on the effects of stress on corrosion pitting growth in aluminium alloys [14]. It was found that corrosion pit growth could be affected

by time, stress amplitude and frequency in a fatigue environment. The methodology of Ishihara et al. [14] was applied to Ni-based superalloys by Chan et al. [15]. They considered the point at which fatigue crack growth exceeds corrosion pit growth. However, neither of these

studies consider the effect of hot corrosion on the mate rials stress intensity threshold (kth), the threshold below which cracking does not occur.

Finite element analysis (FEA) is a commonly used method to calculate stresses within complex geometries or multiaxial loading states. This is done by meshing the geometry as a net of elements and nodes. The ele ments can deform as constrained by the material model,whereas the load is transferred from element to element through the node connections. FEA has been widely used to assess the stress in statically and cyclically loaded conditions.

Experimental method C-ring test method

C-ring specimens were manufactured from CMSX-4 bars. Guidelines for the dimensions were taken from ISO 7539-5 [16]. The final dimensions for the speci mens used in this testing are given in Figure 1. C-ring specimens where manufactured with a < 001 > crystal

lographic orientation aligned with the cylinder axis.

For target stress levels at a constant stain, the required displacement of the C-rings were calculated by first calculating the change in diameter (ΔD) required to achieve a given stress (Equation (1)).

Equation (1): Change in diameter from ISO 7539-5 [16].

ΔD = aπd2∕4EtZ    (1)

FEA modelling was used to verify the stress calculations. Surrogate data from Siebörger et al. [17] for CMSX-4 provided the Young’s modulus (E) for Equation (2) and the monotonic material properties used in FEA mod elling. The final stressed diameters (Df ) were calculated

using Equation (2):  Df = DAV − ΔD  (2)

The C-rings were clamped to the final diameter (Df ) using A2 grade stainless steel M5 nuts, bolts and washers, and measured using a digital micrometre with a resolution of 1  μm (and accuracy of 2  μm). An average of five readings was used to determine the initial external diameter (DAV) from which the final stressed diameter is calculated.  These are given in Table 2.