Skip to main content

Full text of "NASA Technical Reports Server (NTRS) 20000033269: Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments"

See other formats



T vVb - 6 ( 


A Paper Entitled 


Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 


Submitted for presentation in the 


International Joint Tribology Conference 

Sponsored by ASME International and STLE 
October 1-4, 2000 
Seattle, Washington 


And for publication in the ASME Journal of Tribology 


By 


Nagaraj K. Arakere 

Associate Professor 
Mechanical Engineering Department 
University of Florida 
Gainesville, FL 3261 1-6300 
(352) 392-0856: Tel; (352) 392-1071: Fax 


and 

Gregory Swanson 

NASA George C. Marshall Space Flight Center 
ED22/Strength Analysis Group 
MSFC, Alabama-35812 


*■£ * Niagara] K . Arakere and Gregory Swanson 

K ' Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 


Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 


ABSTRACT 

Single crystal nickel base superalloy turbine blades are being utilized in rocket engine turbopumps 
and turbine engines because of their superior creep, stress rupture, melt resistance and 
thermomechanical fatigue capabilities over polycrystalline alloys. Currently the most widely used 
single crystal nickel base turbine blade superalloys are PWA 1480/1493 and PWA 1484. These 
alloys play an important role in commercial, military and space propulsion systems. High Cycle 
Fatigue (HCF) induced failures in aircraft gas turbine and rocket engine turbopump blades is a 
pervasive problem. Blade attachment regions are prone to fretting fatigue failures. Single crystal 
nickel base superalloy turbine blades are especially prone to fretting damage because the subsurface 
shear stresses induced by fretting action at the attachment regions can result in crystallographic 
initiation and crack growth along octahedral planes. Furthermore, crystallographic crack growth on 
octahedral planes under fretting induced mixed mode loading can be an order of magnitude faster 
than under pure mode I loading. 

This paper presents contact stress evaluation in the attachment region for single crystal turbine 
blades used in the NASA alternate Advanced High Pressure Fuel Turbo Pump (HPFTP/AT) for the 
Space Shuttle Main Engine (SSME). Single crystal materials have highly orthotropic properties 
making the position of the crystal lattice relative to the part geometry a significant factor in the 
overall analysis. Blades and the attachment region are modeled using a large-scale 3D finite element 
(FE) model capable of accounting for contact friction, material orthotrophy, and variation in 
primary and secondary crystal orientation. Contact stress analysis in the blade attachment regions is 
presented as a function of coefficient of friction and primary and secondary crystal orientation. 
Stress results are used to discuss fretting fatigue failure analysis of SSME blades. Attachment 
stresses are seen to reach peak values at locations where fretting cracks have been observed. 
Fretting stresses at the attachment region are seen to vary significantly as a function of crystal 
orientation. Attempts to adapt techniques used for estimating fatigue life in the airfoil region, for 
life calculations in the attachment region, are presented. An effective model for predicting 
crystallographic crack initiation under mixed mode loading is required for life prediction under 
fretting action. 


Key Words 

Fretting, Fatigue, High Cycle Fatigue (HCF), Single crystals. Crystal orientation, Nickel-base 
superalloys, Face Centered Cubic (FCC), Turbine blades, Blade attachments 


Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

Introduction 

Fretting fatigue failure of mechanical components has gradually come to be recognized as a failure 
mode of major importance. The presence of fretting in conjunction with a mean stress in the body of 
a component can lead to a marked reduction in high cycle fatigue (HCF) life, sometimes by a factor 
as great as 10 [1]. Fretting occurs when assemblies of components such as blade and disk 
attachment surfaces, bolt flanges, snap fit areas, and other clamped members are subjected to 
vibration, resulting in contact damage. The combined effects of corrosion, wear, and fatigue 
phenomena at the fretting contact facilitates the initiation and subsequent growth of cracks, 
ultimately leading to failure. Single crystal nickel base superalloy turbine blades are especially 
prone to fretting damage because the subsurface shear stresses induced by fretting action at the 
attachment regions can result in crystallographic initiation and crack growth along octahedral 
planes. 

Single crystal materials have highly orthotropic properties making the position of the crystal lattice 
relative to the part geometry a significant factor in the overall analysis. Study of failure modes of 
single crystal turbine blades therefore has to account for material orthotrophy and variations in 
crystal orientation. HCF induced failures in aircraft gas-turbine engines are a pervasive problem 
affecting a wide range of components and materials. HCF accounts for 24% of component failures 
in gas turbine aircraft engines, of which blade failures account for 40% [2], Estimation of blade 
fatigue life, therefore, represents an important aspect of durability assessment. The modified 
Goodman approach currently used for life assessment does not address important factors that affect 
High Cycle Fatigue (HCF) such as fretting and/or galling surface damage, and interaction with Low 
Cycle Fatigue (LCF) [2], Designing to resist fretting fatigue has proven to be a difficult task. The 
fretting fatigue process is affected by a large number of variables, perhaps by as many as fifty [3], 
For polycrystalline materials the critical primary variables are thought to be the coefficient of 
friction, the slip amplitude, and the contact pressure at the fretting interface [1]. However, for single 
crystal nickel superalloys, studying the impact of fretting requires consideration of additional 
variables. The crack initiated by fretting action can be crystallographic or noncrystallographic 
depending on the interaction between the point-source defect species (carbides, micropores and 
eutectics) with the effects of environment, temperature and stress [4]. The impact of fretting on 
relationships between macroscopic fracture details and the parameters that are used to describe 
fatigue crack growth is essential. Understanding the effects of fretting damage on HCF life of single 
crystal nickel superalloy components is critical to lay a foundation for a mechanistic based life 
prediction system. 

Turbine blades and vanes, used in aircraft and rocket engines, are probably the most demanding 
structural applications for high temperature materials [5] due to the combination of high operating 
temperature, corrosive environment, high monotonic and cyclic stresses, long expected component 
lifetimes, and the enormous consequence of structural failure. Directionally solidified (DS) 
columnar-grained and single crystal superalloys have the highest elevated-temperature capabilities 
of any superalloys. As the understanding of alloying and microstructural control (i.e., directional 
solidification) has improved, the maximum use temperatures of Ni-base superalloys has increased to 
levels in excess of 90% of their melting temperatures (T/T m = 0.9). Directional solidification is used 
to produce a single crystal [6] with the <00 1> low modulus orientation parallel to the growth 
direction. The secondary direction normal to the growth direction, which is typically referenced to 
a line parallel to the blade attachment, is random if a grain selector is used to form the single crystal. 
Both the primary and secondary crystal orientations, though, can be selected if seeds are used to 


3 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

generate the single crystal. However, in most cases, grain selectors are used to produce the desired 
<001> growth direction. In this case, the secondary orientations of the single crystal components 
are determined but not controlled. Initially, control of the secondary orientation was not considered 
necessary [7]. However, recent reviews of space shuttle main engine (SSME) turbine blade lifetime 
data has indicated that secondary orientation has a significant impact on high cycle fatigue 
resistance [8,9,10], 

Currently the most widely used single crystal nickel base turbine blade superalloys are PWA 1480, 
PWA 1484, RENE’ N-5 and CMSX-4. PWA 1480 will be used in the NASA SSME alternate 
turbopump. These alloys play an important role in commercial, military and space propulsion 
systems [4]. Military gas turbine mission profiles are characterized by multiple throttle excursions 
associated with maneuvers such as climb, intercept and air-to-air combat. This mission shifts 
attention somewhat to fatigue and fracture considerations associated with areas below the blade 
platform, which contain various stress risers in the form of buttresses and attachments [4], The gas 
turbine industry is moving toward a design philosophy that stresses damage tolerance, structural 
integrity and threshold based designs, which has prompted the study of micromechanics of fretting 
fatigue and other events that affect HCF life. 

Numerous analytical and experimental studies have been conducted on fretting fatigue damage in 
polycrystalline alloys. Some representative examples are by Hills and Newell [1], Giannokopoulos 
and Suresh [11], Swolwinsky and Farris [12], Attia and Waterhouse [13], Hoeppner [14], Vingsbo 
and Soderberg [15], and Ruiz, et al [16], Studies on mechanics of fretting fatigue in single crystal 
superalloys are almost nonexistent. This paper presents an investigation of fretting stresses in single 
crystal superalloy turbine blades in the blade-disk attachment regions. Large-scale finite element 
models of are used to predict stresses in the airfoil and blade attachment regions. Several fatigue 
damage parameters based on shear and normal stresses acting on various slip planes for FCC 
crystals are evaluated. Blade stress response is presented as a function of primary and secondary 
crystallographic orientation. 

Deformation Mechanisms in FCC Single Crystals 

Nickel based single-crystal material such as PWA1480 and PWA1484 are precipitation 
strengthened cast single grain superalloys based on the Ni-Cr-Al system. The microstructure 
consists of approximately 60% to 70% by volume of y’ precipitates in a y matrix. The y’ precipitate, 
based on the intermetallic compound Ni 3 Al, is the strengthening phase in nickel-base superalloys 
and is a Face Centered Cubic (FCC) structure. The / precipitate is suspended within the y matrix 
also has a FCC structure and is comprised of nickel with cobalt, chromium, tungsten and tantalum 
in solution. 

Deformation mechanisms in single crystals are primarily dependent on microstructure, orientation, 
temperature, and crystal structure. The operation of structures at high temperature places additional 
materials constraints on the design that are not required for systems that operate at or near room 
temperatures. In general, materials become weaker with increasing temperature due to thermally 
activated processes, such as multiple slip and cross-slip. At temperatures in excess of approximately 
half the homologous temperature (the ratio of the test temperature to the melting point, = T/T m ), 
diffusion controlled processes (e.g., recovery, recrystallization, dislocation climb and grain growth) 
become important, which results in further reductions in strength. Slip in metal crystals often occurs 


4 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

on planes of high atomic density in closely packed directions. The 4 octahedral planes 
corresponding to the high-density planes in the FCC crystal have 3 primary slip directions (easy- 
slip) resulting in 12 independent primary <1 10> {111} slip systems. The 4 octahedral slip planes 
also have 3 secondary slip directions resulting in 12 independent secondary <11 2> {111} slip 
systems. In addition, the 3 cube slip planes have 2 slip directions resulting in 6 independent <1 10> 
{100} cube slip systems. Thus there are 12 primary and 12 secondary slip systems associated with 
the 4 octahedral planes and 6 cube slip systems with the 3 cube planes, for a total of 30 slip systems 
[10]. At high temperatures, slip has been observed in non-close-pack directions on the octahedral 
plane, and on the cube plane, in FCC crystals. Table 1 shows the 30 possible slip systems in a FCC 
crystal [17]. Elastic response of FCC crystals is obtained by expressing Flooke’s law for materials 
with cubic symmetry. 


Shear stresses associated with the 30 slip systems are denoted by T 1 , T 2 ,..., t 30 . The shear stresses on 
the 24 octahedral slip systems are given by [17], 


V' 


" 1 

0 

-1 

1 

0 

-f 


T 13 ' 



2 

-1 

1 

-2 

1 ’ 

T 2 


0 

-1 

1 

-1 

1 

0 


T 14 


2 

-1 

-1 

1 

1 

-2 

T 3 


1 

-1 

0 

0 

1 

-1 


T 15 


-1 

-1 

2 

-2 

1 

1 

T 4 


-1 

0 

1 

1 

0 

-1 


°xx 


T 16 


-1 

2 

-1 

-1 

- 2 

-1 

T 5 


-1 

1 

0 

0 

-1 

-1 


a yy 


T 17 


-1 

-1 

2 

2 

1 

-1 

T 6 

1 

0 

1 

-1 

-1 

”1 

0 




r 18 

1 

2 

-1 

-1 

-i 

1 

2 

T 7 


1 

-1 

0 

0 

-1 

-1 


a xy 

s 

T 19 

3>/2 

-1 

-1 

2 

2 

-1 

1 

T 8 


0 

1 

-1 

-1 

1 

0 


G zx 


T 20 


2 

-1 

-1 

-1 

-1 

-2 

T 9 


l 

0 

-1 

-1 

0 

~1 


a yz 


T 21 


-1 

2 

-1 

-1 

2 

1 

T 10 


0 

-1 

1 

-1 

-I 

0 


T 22 


2 

-1 

-1 

1 

-1 

2 

T U 


-l 

0 

1 

-1 

0 

-1 


to 

OJ 


-1 

2 

-1 

1 

2 

-1 

T 12 


_-i 

1 

0 

0 

1 

-1 


t 2 \ 


-1 

-1 

2 

-2 

-1 

-1_ 


And the shear stresses on the 6 cube slip systems are 


T 25 ' 


"0 

0 

0 

1 

1 

0“ 



T 26 


0 

0 

0 

1 

-1 

0 


a » 

T 27 

1 

0 

0 

0 

1 

0 

1 


°zz 

to 

OO 


0 

0 

0 

1 

0 

-1 


G xy 

T 29 


0 

0 

0 

0 

1 

1 



k 


_o 

0 

0 

0 

-1 

1 _ 


° yz . 


( 2 ) 


Shear strains (engineering) on the 30 slip systems are calculated using similar kinematic relations. 

The stresses (a xx , cr yy , , a yz ) must be expressed in the material or the crystal coordinate system, 

i.e., the x, y, and z coordinate axes are parallel to the edges of the FCC crystal. 

Fretting Fatigue Crack Initiation and Crack Growth in Nickel-Base Superalloys 

Crack initiation in Ni-base superalloys fall into two broad classifications, crystallographic and 
noncrystallographic. The interaction between the effects of environment, temperature and stress 


5 



'Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

intensity determines which point-source defect species initiates a crystallographic or 
noncrystallographic fatigue crack [4]. At low temperature (<427 C) and stress conditions 
crystallographic initiation appears to be the most prevalent mode. At moderately high temperature 
(above 593C) initiation at TaC carbides predominates over occasional (111) eutectic origins. The 
subsurface shear stresses induced by fretting action can result in crystallographic initiation of 
failure, as seen in Fig. 1, which shows a subsurface fatigue crack emanating from a carbide (TaC) in 
a turbine blade attachment (PWA1422, a columnar grain Ni-base superalloy) and propagating along 
octahedral (111) planes [4], Fretting fatigue at low slip amplitudes that induces little or no surface 
damage can result in greatly reduced fatigue life with accelerated subsurface crystallographic crack 
initiation, akin to subsurface shear stress induced rolling bearing fatigue. An example of this 
phenomenon is seen in Fig. 2, which details the underside view of a single crystal Ni (PWA1480) 
turbine blade platform tip fretting failure at the contact where a centrifugal damper impinged the 
blade platform [18]. The damper (or fret pin) was subjected to an alternating compressive load 
generating subsurface shear stresses. The fretting fatigue crack initiation was clearly subsurface, 
propagating along intersecting (111) crystallographic planes until breaking out to the surface. The 
piece was subsequently ejected. An identical scenario occurred with the impinging damper. The 
damper was quite small and cast in a very coarse grained form such that the contact point was a 
single large grain. It too exhibited a subsurface crystallographic shear fatigue crack 
initiation/propagation on multiple intersecting octahedral planes leading to a pyramidal "hole" in the 
component as the "chip" was ejected. 

The fatigue crack growth (FCG) behavior of single crystal nickel superalloys is governed by a 
complex interaction between the operative deformation mechanism, stress intensity, and 
environmental conditions. The FCG behavior is determined by the operative microscopic fracture 
mode. As a result of the two phase microstructure present in single crystal nickel alloys a complex 
set of fracture mode exists based on the dislocation motion in the matrix ( 7 ) and precipitate phase 
(/). Telesman and Ghosn [19] have observed the transition of fracture mode as a function of stress 
intensity (K) in PWA1480 at room temperature. They also noted that, K rss , the resolved shear stress 
intensity parameter on the 12 primary slip planes predicted the crack zigzag behavior caused by 
fracture mode transitions. Deluca and Cowles [20] have observed the fracture mode transition that is 
environmentally dependent (presence of high-pressure hydrogen). The resolved shear stresses and 
not the maximum principal stresses govern failure on the (111) octahedral slip planes. However 
stresses normal to the plane (mode I component) are thought to play some role in the failure 
process. Crystallographic crack growth along (111) planes under mixed mode loading can be an 
order of magnitude faster than under pure mode I loading [21]. These results have important 
implications on fretting fatigue. Since fretting action results in mixed mode loading this can result 
in crystallographic initiation along (111) slip planes (Fig. 1) and rapid crack growth under loads 
lower than that expected under mode I loading. 

Effect of Crystal Orientation on Fatigue Life 

Single crystal materials have highly orthotropic properties making the position of the crystal lattice 
relative to the part geometry a significant factor in the overall analysis. For example, the elastic or 
Young’s modulus (E) is a strong function of crystal orientation over the standard stereographic 
triangle, given by Eq. (3) [22]. 

E' 1 = Sn - [2(Sn - S, 2 ) - S 44] [cos 2 <t>(sin 2 4 - sin 2 0 cos 2 <J> cos 2 0)] (3) 


6 


Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 


Where 0 is the angle between the growth direction and <001> and <|> is the angle between the <001 > 

- <110> boundary of the triangle. The terms Sn, S 12 and S 44 are the elastic compliances. In 
general, the elastic properties are insensitive to composition and values for pure Ni are a good 
approximation of the values for even the most highly alloyed Ni-base superalloys. 

Strain controlled LCF tests conducted at 1200 F in air for PWA1480/1493 uniaxial smooth 
specimens, for four different orientations, is shown in Table 2 [9]. The four specimen orientations 
are <00 1> (5 data points), < 1 1 1 > (7 data points), <213> (4 data points), and <01 1> (3 data points), 
for a total of 19 data points. Figure 3 shows the plot of strain range vs. Cycles to failure. A wide 
scatter is observed in the data with poor correlation for a power law fit. Table 3 shows the fatigue 
parameters such as AT max , Acr max , A£, Ay , etc computed on the 30 slip systems. The procedure for 
obtaining values shown in Table 3 is explained in greater detail by Arakere and Swanson [9]. Other 
fatigue parameters used in polycrystalline material were also evaluated in reference [9], besides 
AT max . Figure 4 shows that the maximum shear stress amplitude AT max plotted against cycles to 
failure, N, has a good correlation for a power law fit. The power law curve fit is shown by Eq. (4). 

Ax max = 397,758 N' 01598 (4) 

The correlation for [AT max ] would be better if some of the high stress data points are corrected for 
inelastic effects. It must be noted that Eq. (4) is only valid for PWA1480/1493 material at 1200F. 
Since the deformation mechanisms in single crystals are controlled by the propagation of 
dislocations driven by shear the [AT max ] might indeed be a good fatigue failure parameter to use for 
crystallographic cracking. However, this parameter must be verified for a wider range of R-values 
and specimen orientations, and also at other temperatures and environmental conditions. Equation 
(4) will be used to calculate fatigue life at a critical blade tip location for the SSME turbine blade. 

Fatigue Failure in SSME HPFTP/AT Single Crystal Turbine Blades 

Turbine blades used in the alternate high-pressure fuel turbopump (HPFTP/AT) for the NASA 
SSME are fabricated from single crystal nickel PWA1480/1493 material. Many of these blades 
have failed during testing due to fretting/galling fatigue cracks in the attachment regions and the 
initiation and propagation of fatigue cracks from an area of high concentrated stress at the blade tip 
leading edge. The subsequent investigation into the blade tip cracking failure provided insight into 
areas where the robustness of the design could be improved to reduce the potential for failure. 
During the course of the investigation an interesting development was brought to light. When the 
size of the fatigue cracks for the population of blades inspected was compared with the secondary 
crystallographic orientation (3 a definite relationship was apparent as shown in Fig. 5 [8, 23], 
Secondary orientation does appear to have considerable influence over whether a crack will initiate 
and arrest or continue to grow until failure of the blade airfoil occurs. Figure 5 reveals that for (3 = 
45+/- 15 degrees tip cracks arrested after some growth or did not initiate at all. This suggests that 
perhaps there are preferential [3 orientations for which crack growth is minimized at the blade tip. 

In an attempt to understand the effect of crystal orientation on blade stress response at the blade tip 
and in the attachment regions a detailed three-dimensional finite element (FE) model capable of 
accounting for primary and secondary crystal orientation variation was constructed [10, 24], 


7 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

Description of the Blade Finite Element (FE) Model 

The blade FE model is described in detail by the present authors in reference [9], which deals with 
the effects of crystal orientation on blade fatigue life. A brief description of the model is repeated 
here for completeness. The Alternate High Pressure Fuel Turbo Pump (HPFTP/AT) first stage blade 
ANSYS FE model was constructed from a large 3D cyclic symmetry model that includes the first 
and second stage blades and retainers, interstage spacer, disk and shaft, and the disk covers [10]. 
The FE models are geometrically nonlinear due to the contact surfaces between the separate 
components. The element type used for the blade material is the ANSYS SOLID45, an 8-noded 3D 
solid isoparametric element. Anisotropic material properties are allowed with this element type. 
The FE model of the first stage turbine blade with the frequently observed blade tip cracking 
problem is shown in Fig. 6. Figure 6 also shows the material coordinate system, which is 
referenced, to the blade casting coordinate system. 

To examine a wide range of primary and secondary variations in crystal orientation, 297 material 
coordinate systems were generated for this study. The primary and secondary crystallographic 
orientation is shown in Fig. 7 [8]. Figure 8 shows the distribution of the 297 different material 
coordinate systems within the allowed 15-degree maximum deviation from the casting axis. The 
secondary repeats after 90 degrees, so only 0 to 80 degrees needed to be modeled. The two angles, 
A and T, locate the primary material axis relative to the casting axis, and the third angle, [3, is the 
clocking of the secondary material axis about the primary material axis. 

The loading conditions represent full power mainstage operation of the Space Shuttle Main Engine, 
referred to as 109% RPL SL (Rated Power Level Service Life). The shaft speed is 37,355 rpm, the 
airfoil temperature is approximately 1200 F, forces representing the blade damper radial sling load 
are applied to the blade platform, and aerodynamic pressures are applied to the blade surfaces and 
internal core. 

The connection between the blade and disk are modeled with ANSYS COMBIN40 elements, for 
cases where the coefficient of friction between the blade-disk attachments was ignored. These 
elements have one degree of freedom at each node. The nodal motion in that degree of freedom sets 
the separation or contact of these elements only. This element does not have the capability for 
friction tangent to the contact surface. 

To account for coefficient of friction at the blade attachment regions, the COMBIN40 element was 
replaced with CONTAC52 type elements. The CONTAC52 element has 3 degrees of freedom at 
each node. Coefficient of friction at the contact is allowed, thus developing tangential traction 
forces. The relative position and motion of the end nodes are used to define displacement and 
reaction force magnitude and direction at the contact. For FE models including friction at the 
attachment regions, 

Post processing of the 297 FE output files represented a considerable amount of work. Two 
FORTRAN programs, developed at MSFC, were employed for this effort. The first strips the 
element results from the coded binary output files and places them into ASCII text files. The 
second program processes the ASCII files to calculate averaged nodal results, the resolved shear 
stresses and strains and the normal stresses and strains in the single crystal material coordinate 
system. Fatigue parameters chosen for study are then calculated and sorted based on user set 


8 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

criteria. Postprocessing of the friction data was performed by constructing a local coordinate 
system with one axis in the tangential direction of contact slip on the blade attachment (Fig. 14). 
An ANSYS batch-postprocessing routine was written to output tabulated results for selected nodes 
in the attachment. 

Effect of Crystal Orientation on Blade Tip Stress Response 

Effect of crystal orientation on stress response at the blade tip critical point is discussed first, 
followed by a discussion of stresses in the blade attachment region. The crack location and 
orientation at the critical blade tip location is shown in Fig. 5. Stress response at the blade tip as a 
function of crystal orientation is discussed in greater detail in reference [9]. 

Fatigue life estimation of single crystal turbine blade components represents an important aspect of 
durability assessment. Towards identifying a fatigue damage parameter for single crystal blade 
material, several fatigue failure criteria used for polycrystalline material subjected to multiaxial 
states of fatigue stress were evaluated for single crystal material by Arakere and Swanson [9]. They 
are outlined here briefly to facilitate comparisons with fatigue in contact locations. 

Turbine blade material is subjected to large mean stresses from the centrifugal stress field. High 
frequency alternating fatigue stresses are a function of the vibratory characteristics of the blade. 
Kandil, et al [25] presented a shear and normal strain based model, shown in Eq. 5, based on the 
critical plane approach, which postulates that cracks initiate and grow on certain planes and that the 
normal strains to those planes assist in the fatigue crack growth process. In Eq. (5) y nia x is the max 
shear strain on the critical plane, e n the normal strain on the same plane, S is a constant, and N is the 
cycles to initiation. 


7 m ax+5£„=/W (5) 

Socie, et al [26] presented a modified version of this theory shown in Eq. (6), to include mean stress 
effects. Here the maximum shear strain amplitude (A)) is modified by the normal strain amplitude 
(Ae) and the mean stress normal to the maximum shear strain amplitude (c n o)- These observations 
are similar to FCG on octahedral planes for single crystals. 


A r t A£ n 
2 2 



= f(N) 


( 6 ) 


Fatemi and Socie [27] have presented an alternate shear based model for multiaxial mean-stress 
loadings that exhibit substantial out-of-phase hardening, shown in Eq. (7). This model indicates that 
no shear direction crack growth occurs if there is no shear alternation. 


Ar <7 max 

+ ~) = f(N) (7) 

2 Gy 

Smith, Watson, and Topper [28] proposed a uniaxial parameter to account for mean stress effects 
which was modified for multiaxial loading, shown in Eq. (8), by Bannantine and Socie [29]. Here 


9 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

the maximum principal strain amplitude is modified by the maximum stress in the direction of 
maximum principal strain amplitude that occurs over one cycle. 

Ac 

T (cr max ) = f (N) (8) 

Equation 8 has been used successfully for prediction of crack initiation in fretting fatigue, for 
aluminum and titanium polycrystalline alloys [12]. The polycrystalline failure parameters based on 
equations (5-8) was applied for single crystal uniaxial LCF test data, shown in Table 2, by 
evaluating the fatigue damage parameters on all of the 30 slip systems [9, 10]. However, the most 
effective fatigue damage parameter was found to be maximum shear stress amplitude AT max shown 
in Fig. 4. 

Variation of crystal orientation on stress response at the blade tip critical point prone to cracking 
(tip point on inside radius) was examined by analyzing the results from the 297 FE model runs. The 
FE node at the critical point was isolated and critical failure parameter value (Ax max ) computed on 
the 30 slip systems. A contour plot of At max was generated as a function of primary and secondary 
orientation, shown in Fig. 9. The contour plot clearly shows a minimum value for Ax max for 
secondary orientation of P = 50 deg and primary orientation designated by cases 5 and 20. Case 5 
corresponds to a primary orientation of A = 0 deg and T = 7.5 degree. Case 20 corresponds to a 
primary orientation of A = 5.74 deg and T = 13.86 degree. Using the fatigue life equation based on 
the Atmax curve fit of LCF test data (Eq. 4) we can obtain a contour plot of dimensionless life at the 
critical point as a function of primary and secondary orientation, as shown in Fig. 10. The maximum 
life is again obtained for p = 50 deg, and A = 0 deg and T = 7.5 deg, and A = 5.74 deg and T = 1 3.86 
deg. The optimum value of secondary orientation P = 50 deg. corresponds very closely to the 
optimum value of p indicated in Fig. 5. This demonstrates that control of secondary and primary 
crystallographic orientation has the potential to significantly increase a component's resistance to 
fatigue crack growth without adding additional weight (for the airfoil region). 

Stress Response in the Attachment Region (Frictionless Contact) 

The 297 FE models representing 33 variations (cases) in primary crystal orientation and 9 variations 
in the secondary orientation were run for frictionless contacts at the blade attachments. Blade 
loading conditions were kept constant while the crystal orientation was varied. Figure 1 1 shows 
representative stress plots for the blade attachment region. The critical point with the most severe 
contact stresses is shown in Fig. 12. Fretting damage in the attachment regions is highlighted in 
Figs. 13-14. Effects of variation of primary and secondary orientation on stress response are studied 
for this critical region. Figures 15-16 show contour plots of At max and t max * Ay/2, as a function of 
primary and secondary orientation. The crystal orientation does not appear to have any discemable 
correlation at the contact point for these parameters. Figure 17 shows a contour plot of 
(G max Ae/2, Eq. 8) as a function of primary and secondary orientation. The secondary orientation 
does appear to have an optimum value of about 55 degrees. Szolwinski and Farris [12] have used 
Eq. (8) successfully as a fretting damage parameter for polycrystalline alloys. For single crystal 
material, however, the fretting damage parameters need to be evaluated in greater detail. A simpler 
geometry such as a cylindrical fret pin made of polycrystalline material contacting a flat plate made 
of single crystal material with a specific crystal orientation would be amenable to an analytical 


10 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

solution. Surface and subsurface stresses at the contact can then be examined in detail, and fretting 
damage parameters evaluated, as a function of crystal orientation, coefficient of friction, slip 
amplitude, loads, etc. 

Figures 1 and 2 illustrate that subsurface shear stresses from fretting action, at low or even zero slip 
amplitude, can induce subsurface crystallographic crack initiation, with stage I-fatigue cracks 
propagating outward to the surface. No crack involvement of the fretted surface occurs and 
negligible surface damage is noted at the contact surfaces. Clearly, an effective model for predicting 
crystallographic crack initiation under mixed mode loading is required for life prediction under 
fretting action. 

This study shows that stress at the attachment region varied significantly as a function of crystal 
orientation alone, from x max = 60 ksi to x max = 130 ksi. This indicates that by selecting the appropriate 
crystal orientation, for specified geometry and loading, it may be possible to optimize blade 
structural integrity. 

Stress Response in the Attachment Region Including Contact Friction 

The presence of friction results in the development of traction forces at the fretting contact. 
Evaluation of the contact forces and the resulting contact stresses at the interface require careful 
analysis of the interaction between the contacting surfaces and applied loads. To study the effect of 
variation of friction coefficient at the contact, and variation of primary and secondary crystal 
orientations, 18 FE model runs were completed. The 18 FE models comprised of 3 primary 
orientations (cases 0, 5 and 30), 2 secondary orientations (P = 0 and 50 deg.), and 3 values for 
coefficient of friction ((l = 0, 0.3, 0.7). Figure 12 shows representative contact stress results in the 
critical attachment region. Blade loading is kept constant for all 18 cases. Figure 18 shows contour 
plots of tangential normal stress, ax, at the contact surface. The X coordinate is in the direction of 
slip, as shown in Fig. 12. The tangential normal stress is of practical interest because cracks are 
thought to initiate at locations where ax reaches a maximum tensile value. The maximum value of 
tangential normal stress, ax, reached in the contact zone is listed in Table 4 for all the 18 cases. It is 
seen that <J\ increases with jx. ax is also seen to vary considerably with variations in primary and 
secondary orientation. Since the component stiffness varies with crystal orientation the stress 
distribution is also expected to vary, under constant loading. Minimum values of G\ are reached for 
case 30 and p=0 [(84.3 ksi, (1=0), (89.9 ksi, (1=0.3), (97.7 ksi, (1=0.7)] indicating that perhaps key 
design parameters can be optimized for specific blade geometry and loading. Additional fretting 
damage parameters need to be examined, for design optimization possibilities. 

The frictional contact study showed that both a x and x max stresses at the attachment region varied by 
about 20%, as a function of crystal orientation alone. For instance, the ax variations (Table 4) as a 
function of crystal orientation alone are 84.3 - 108.6 ksi for (I = 0, 89.9 - 1 14.4 ksi for (i = 0.3, and 
97.7 - 122.6 ksi for (t = 0.7. The variations would have been larger if peak variations due to crystal 
orientation were realized. It appears that if case 25 with P=30° was compared with case 17 at P=0° 
or, P=20° stress variations due to crystal orientation would have been considerably larger than 20%. 

Figures 13 and 14 show representative fretting/galling induced cracks in the blade attachment 
region for HPFTP/AT first stage blades. Region shown in Fig. 13 is on the suction side, trailing 


11 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

edge of the blade. Several arrest marks are also visible. The blade crystal orientation is A = -6.7°, y 
= 11.3°, |3 =4.2°. Figure 14 shows fretting/galling induced cracking showing multiple origins with 
stage II cracks. The crystal orientation for this blade is A = -2°, y = 3°, (3 =7°. Many blades in 
different engine units exhibited similar blade attachment cracks. In the blade casting process, the 
secondary crystal orientation is not controlled, while the primary crystal orientation is controlled to 
within 15°. A systematic investigation of severity of fretting/galling induced attachment cracks, 
similar to study shown in Fig. 5 for blade tip cracking, has to be done to discern relationships 
between fretting damage and crystal orientation. Tolerance in clearance between the blade and disk 
attachment surfaces also plays an important role in the contact stress distribution. The analysis 
results presented are for nominal clearances between mating parts. 

Conclusions 

Three dimensional finite element stress analysis of HPFTP/AT SSME single crystal turbine blades 
subjected to rotational, aerodynamic, and thermal loads is presented. Stress response at the blade tip 
and attachment regions are presented as a function of primary and secondary crystal orientation. 297 
FE models are analyzed to study a wide range of variation in crystal orientation. For the blade tip 
location, the maximum shear stress amplitude [Ar max ] on the 30 slip systems was found to be an 
effective fatigue failure criterion, based on the curve fit between AT max and LCF data at 1200F for 
PWA1480/1493. Variation of At max as a function of crystal orientation at the blade tip reveals that 
control of secondary and primary crystallographic orientation has the potential to significantly 
increase a component's resistance to fatigue crack growth without adding additional weight or cost. 
Current seeding techniques used in the single crystal casting process can readily achieve the degrees 
of control in primary and secondary crystallographic orientation required. Fretting contact stresses 
are presented for the attachment region. Attachment stresses are seen to reach peak values at 
locations where fretting cracks have been observed. For frictionless contact, stress at the attachment 
region varied significantly as a function of crystal orientation alone, from T max =60 ksi to T max = 130 
ksi, under constant loading conditions. Attachment stresses varied significantly for models that 
included coefficient of friction at the contact. The tangential normal stress ax increased with 
increasing coefficient of friction at the critical contact location. This might be of practical interest 
since cracks are thought to initiate at locations where ax reaches a maximum tensile value, for 
polycrystalline materials. The parameter Ax max , which showed a definite correlation with crystal 
orientation at the blade tip, appears to have no discemable correlation at the critical contact region. 
The damage parameter a ma x Ae/2 appears to be correlated with the secondary orientation. Fretting 
damage parameters need to be evaluated in greater detail for single crystal materials. 

Subsurface shear stresses induced by fretting action, at low or even zero slip amplitude, can induce 
subsurface crystallographic crack initiation, with no crack involvement of the fretted surface. An 
effective model for predicting crystallographic crack initiation under mixed mode loading is 
required for life prediction under fretting action. This will necessitate specialized testing to account 
for crystal orientation and mixed mode loading effects. One way to conduct these tests is to use 
Brazilian disk specimens as outlined by John, et al [21], to study mixed mode loading in single 
crystals. However, to study crystallographic crack initiation behavior, the disks would have to be 
tested without starter flaws, requiring high compressive loads and potential for failure at the loading 
points. One solution perhaps is to thin the disks at the center which necessitates lower compressive 
loads to initiate a crystallographic crack at the disk center. 


12 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 


Acknowledgements 

The authors would like to thank the NASA/ASEE 1999 Summer Faculty Fellowship Program. The 
support from this program, administered by the University of Alabama in Huntsville, enabled the 
first author to perform part of this work at the NASA Marshall Space Flight Center, Huntsville, AL. 


REFERENCES CITED 

[J] D. A. Hills and D. Nowell, Mechanics of Fretting Fatigue, Kluwer, Deventer, 1 994. 

[2] B. A. Cowles, High cycle fatigue in aircraft gas turbines-an industry perspective, 
International Journal of Fracture, (1996) 1-16. 

[3] J. Dombromirski, Variables of Fretting Process: Are There 50 of them? Standardization of 
Fretting Fatigue Test Methods and Equipment, ASTM, pp. 60-68, 1990. 

[4] Deluca, D., Annis. C, "Fatigue in Single Crystal Nickel Superalloys"; Office of Naval 
Research, Department of the Navy FR23800, August 1995. 

[5] C.T. Sims, “Superalloys: Genesis and Character”, Superalloys - II, Eds., C.T. Sims, N.S. 
Stoloff and W.C. Hagel, Wiley & Sons, New York, New York, (1987), p. 1. 

[6] F.L. VerSnyder and R.W. Guard, “Directional Grain Structure for High Temperature 
Strength”, Trans. ASM, 52, (1960), p. 485. 

[7] M. Gell and D.N. Duhl, “The Development of Single Crystal Superalloy Turbine Blades”, 
Processing and Properties of Advanced High-Temperature Materials, Eds., S.M. Allen, R.M. 
Pelloux and R. Widmer, ASM, Metals Park, Ohio, (1986), p. 41. 

[8] J. Moroso, “Effect of Secondary Orientation on Fatigue Crack Growth in Single Crystal 
Turbine Blades”, M.S. Thesis, Mechanical Engineering Department, University of Florida, 
Gainesville, FL, May 1999. 

[9] N. K. Arakere and G. Swanson, “Effect of Crystal Orientation on Fatigue Failure of Single 
Crystal Nickel Base Turbine Blade Superalloys,” accepted for presentation in the ASME 
IGTI conference (May 8-1 1, 2000, Munich, Germany) and for publication in the ASME 
Journal of Gas Turbines and Power. 

[10] N. K. Arakere and G. Swanson, “Fatigue Failure of Single Crystal Nickel Base Turbine 
Blade Superalloys,” accepted for publication as a NASA Technical Paper. 

[11] Giannokopoulos and S. Suresh, “3D Fretting Stresses,” Acta Materialia, 1999. 

[12] M. P. Szolwinski and T. N. Farris, “Mechanics of Fretting Fatigue Crack Formation,” Wear, 
v 198, 1996, pp. 93-107. 

[13] M. H. Attia and R. B. Waterhouse, Editors, Standardization of Fretting Fatigue Test 
Methods and Equipment, ASTM (04-01 1590-30), STP 1 159, (1992) 

[14] D. W. Hoeppner, Mechanisms of Fretting Fatigue and Their Impact on Test Methods 
Development, Standardization of Fretting Fatigue Test Methods and Equipment, ASTM, pp. 
23-32, 1990. 

[15] O. Vingsbo and D. Soderberg, On fretting maps. Wear, 126 (1988) 131-147. 

[16] C. Ruiz, P. H. B. Boddington and K. C. Chen, An investigation of fatigue and fretting in a 
dovetail joint, Experimental mechanics, 24 (3) (1984) 208-217. 

[17] Stouffer, D. C., and Dame, L. T., “Inelastic Deformation of Metals,” John Wiley & Sons; 
1996. 


13 


Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

[18] Personal Communication with D. P. DeLuca, Pratt & Whitney, (Government Engines and 
Space Propulsion), Mechanics of Materials, West Palm Beach, FL. 

[19] Telesman, J.; Ghosn, L, “The unusual near threshold FCG behavior of a single crystal 
superalloy and the resolved shear stress as the crack driving force,” Engineering Fracture 
Mechanics, Vol. 34, No. 5/6, pp. 1183-1196, 1989. 

[20] Deluca, D. P. and Cowles, B. A., “Fatigue and fracture of single crystal nickel in high 
pressure hydrogen”, Hydrogen Effects on Material Behavior, Ed. By N. R. Moody and A. 
W. Thomson, TMS., Warrendale, PA, 1989. 

[21] John, R., DeLuca, D. P., Nicholas, T., and Porter, J., “Near-threshold crack growth behavior 
of a single crystal Ni-base superalloy subjected to mixed mode loading,” Mixed-Mode Crack 
Behavior, ASTM STP 1359, Editors. K. J. Miller and D. L. McDowell, Paper ID: 5017, 
November 1998. 

[22] M. McLean, “Mechanical Behavior: Superalloys”, Directionally Solidified Materials for 
High Temperature Service, The Metals Society, London, (1983), p. 151. 

[23] Pratt and Whitney, “SSME Alternate Turbopump Development Program HPFTP Critical 
Design Review,” P&W FR24581-1 December 23, 1996. NASA Contract NAS8-36801. 

[24] Sayyah, T., “Alternate Turbopump Development Single Crystal Failure Criterion for High 
Pressure Fuel Turbopump First Stage Blades,” Report No.: 621-025-99-001 , NASA 
Contract NAS 8-40836, May 27, 1999. 

[25] Kandil, F. A., Brown, M. W., and Miller, K. J., “Biaxial low cycle fatigue of 316 stainless 
steel at elevated temperatures,” Book, pp. 203-210, Metals Soc., London, 1982. 

[26] Socie, D. F., Kurath, P., and Koch,J., “A multiaxial fatigue damage parameter,” presented at 
the Second International Symposium on Multiaxial Fatigue, Sheffield, U.K., 1985. 

[27] Fatemi, A, Socie, D, “A Critical Plane Approach to Multiaxial Fatigue Damage Including 
Out-of-Phase Loading,” Fatigue Fracture in Engineering Materials, Yol. 1 1, No. 3, pp. 149- 
165, 1988. 

[28] Smith, K. N., Watson, P., and Topper, T. M., “A stress-strain function for the fatigue of 
metals,” Journal of Materials, Vol. 5, No. 4, pp. 767-778, 1970. 

[29] Banantine, J. A., and Socie, D. F., “Observations of cracking behavior in tension and torsion 
low cycle fatigue,” presented at ASTM Symposium on low cycle fatigue - Directions for the 
future, Philadelphia, 1985. 


14 



Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 


LIST OF FIGURES 

Fig. 1 A subsurface fretting fatigue crack emanating from a carbide in a turbine blade 

attachment (PWA1422) and propagating along octahedral (111) shear planes [4]. 

Fig. 2 Subsurface fretting fatigue crack initiation in a single crystal Ni turbine blade (platform 
tip) [18] 

Fig. 3 Strain range Vs. Cycles to Failure for LCF test data (PWA1493 at 1200F) 

Fig. 4 Shear Stress Amplitude [AT max ] Vs. Cycles to Failure 

Fig. 5 Secondary Crystallographic Orientation, p, Vs Crack Depth for the SSME AHPFTP 1 st 
Stage Turbine Blade [8, 23] 

Fig. 6. First Stage Turbine Blade FE Model and Casting Coordinate System 

Fig. 7 Convention for Defining Crystal Orientation in Turbine Blades [8] 

Fig. 8 33 primary crystallographic orientations (cases) and 9 secondary crystallographic 

orientations ((3 or 0 values), for a total of 297 material orientations. 

Fig. 9 Maximum Shear Stress Amplitude (Ax max , ksi) Contour Plot at the Blade Tip Critical 
Point 

Fig. 10 Normalized HCF life (Contour Plot) at the blade tip Critical Point, as a function of 
primary and secondary orientation 

Fig. 11 Representative stress plots for the single crystal blade attachment region 

Fig. 12 HPFTP/AT first stage blade vonMises stress plot with local zoom in of the suction side 
upper contact region at the blade leading edge and the local coordinate system used for 
the contact results. 

Fig. 13 Fretting/galling induced crack in the contact region (suction side, trailing edge of blade) 
Several arrest marks are visible. Crystal orientation: A = -6.7°, y = 1 1 .3°, p =4.2°. 

Fig. 14 Fretting/galling induced cracking showing multiple origins and stage II cracks (pressure 
side trailing edge location). Crystal orientation: A = -2°, y = 3°, p =7°. 

Fig. 15 Contour plot of max shear stress amplitude, Ax ma x (ksi), at the critical contact location, 
as a function of primary (Case number) and secondary (P or 0) crystallographic 
orientation. 


Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 

Fig. 16 Contour plot of the parameter, T max *(Ay/2), at the critical contact location, as a function 
of primary (Case number) and secondary (P or 0) crystallographic orientation. 

Fig. 17 Contour plot of the parameter, a max *(Ae/2), at the critical contact location, as a function 
of primary (Case number) and secondary (p or 0) crystallographic orientation. 

Fig. 18 Case = 0, P = 0, p = 0, Contour plot of tangential surface stress, G x , in upper lobe, 
suction side, near leading edge. 


LIST OF TABLES 

Table 1 Slip Planes and Slip Directions in a FCC Crystal [17] 

Table 2 Strain controlled LCF test data at 1200 F for 4 specimen orientations 

Table 3 Maximum values of shear stress and shear strain on the slip systems and normal stress 
and strain values on the same planes. 

Table 4 Values of crystal orientation, friction coefficient, and max tangential normal stress ax 
at the critical contact location, for the 1 8 FE model runs 



Slip Plane 

Slip Direction 

Octahedral Slip a/2<l 10>f 1 1 1] 

12 Primary Slip Directions 

[111] 

[1 0 -1] 

[111] 

[0 -i 1] ! 

[111] 

[1 -1 0] 

[-1 1 -1] 

[1 0 -1] 

M 1 -1] 

[110] I 

[-1 1 -1] 

[0 1 1] 

[1 -1 -1] 

[1 1 0] 

[1 -1 -1] 

[0 -1 1] 

[1 -1 -1] 

[1 0 1] 

[-1 -1 1] 

[0 1 1] 

M -1 1] 

[1 0 1] 

[-1 -1 1] 

[1 -1 0] 


Octahedral Slip a/2<l 12>f 1 1 1 


[111] 

[111] 


[111] 


[-1 1 - 1 ] 

[-1 1 - 1 ] 

[-1 1 - 1 ] 

[1 -1 - 1 ] 

[1 -1 - 1 ] 

[1 -1 - 1 ] 


[-1 -1 1 ] 

[-1 -1 1 ] 
[-1 -1 1 ] 


Cube Slip a/2<l 10>{ 100 


[1 0 0] 
[0 1 0] 
[0 1 0] 
[0 0 1] 
[0 0 1] 


12 Secondary Slip Directions 


[-1 2 - 1 ] 


[2 -1 - 1 ] 

[-1 -1 2 ] 


LLAJJ 

[i -i -2] 

[-2 -i n 

[-1 1 - 2 ] 

[2 1 1 ] 

[-1 ~2 1 ] 


[-2 1 - 1 ] 

[1 -2 -1] 

[1 1 2 ] 


6 Cube Slip Directions 


[0 1 1] 


[0 1 -i] 

[1 o n 

[1 0 -i] 

[1 1 0] 

[-1 1 0 ] 


Table 1 Slip Planes and Slip Directions in a FCC Crystal [17] 











































Specimen 

Orientation 

Max 

Test Strain 

Min 

Test Strain 

R 

Ratio 

Strain 

Range 

Cycles 

to 

Failure 

<001 > 

.01509 

.00014 

0.01 

.01495 

1326 

<00 1> 

.0174 

.0027 

0.16 

0.0147 

1593 

<001 > 


.0002 

0.02 

0.011 

4414 

<00 1> 

.01202 

.00008 

0.01 

0.0119 

5673 1 

<001 > 

.00891 

.00018 

0.02 

.00873 

29516 

<1 1 1> 

.01219 

-0.006 

-0.49 

.01819 

26 

<11 1> 

.0096 

.0015 

0.16 

0.0081 

843 

<1 1 1> 

.00809 

.00008 

0.01 

.00801 

1016 

<1 1 1 > 

.006 

0.0 

0.0 

0.006 

3410 

<1 1 1 > 

.00291 

-0.00284 

-0.98 

.00575 

7101 

<11 1> 

.00591 

.00015 

0.03 

.00576 

7356 

<1 1 1> 

.01205 

0.00625 

0.52 

0.0058 

7904 

■ESEaBi 

.01212 

0.0 

0.0 

.01212 

79 


■ 

.00013 

0.02 

.00782 

4175 



.00005 

0.01 

.00596 

34676 

BiH 

.006 

0.0 

0.0 

0.006 

114789 

<01 1> 

.0092 

.0004 

0.04 

0.0088 

2672 

<01 1 > 

.00896 


0.01 

.00883 

7532 

<01 1 > 

.00695 


0.03 

.00676 

30220 


Table 2 Strain controlled LCF test data at 1200 F for 4 specimen orientations 




























g <D 
j=2 3 

> 1 
,g .c 
c3 


ft CA 

c c 

2 S 
S £ 
« ‘3 

o o 
Cu Cl 
«a So 

5 .5 

6 £ 


3 § 

c e 
2 ’3 

b j3 
V5 V 

ft to 

<U <D 
w +- » 

1 § 

e e 

<r> *3 

O (D 
Ch ex 
«« CO 

X C 

«j -5 

£ £ 


0,0 0,0 

lb <— I 4— I 4-4 

00 CO CO CO 

£ £ £ £ 

<D 4> O O 

W W ^ 4— 1 

{/) W (/5 Cfl 
>> ^ 

CO CO CO CO 

CW Oh Oh 

CO CO CO CO 

O o O O 
co m m co 


o o 
e c 
'Si £ 
b b 

CO CO 

& £ 

<D O 

■s -s 

S 3 

s s 

II II 

1 1 

5 s 


o o 

co co 
CO co 
<L> P 

b b 

co CO 

£ £ 

o <d 

•s -s 
3 3 
S S 

ii ii 

s I 
J ^ 


0 1 

*0 O M 

£ ^ « 
W fe 

cn ^ m 'S 
oj ck — r- , 7 : 

£23 8 § 

26 

843 

1016 

3410 

7101 

7356 

7904 

un %c 

S ^ ^ 

04 CS ^ 
ocn ^ 
on g 

04 O- g 

3 >5J 

< °* 

7.68E+04 

7.54E+04 

5.65E+04 

6.17E+04 

4.57E+04 

237E+05 

1.05E+05 

1.05E+05 

7.84E+04 

7.50E+04 

7.50E+04 

8.20E+04 

1.30E+05 

8.46E+04 

6.50E+04 

6.50E+04 

n n n 

?? ? 
IdU Id 
m o 0- 

'sD O- 04 

i * 

2 w § o o 
r- on 2 
m ^ 

-7.80E+04 

1.96E+04 

1045 

0 

-3.70E+04 

1959 

7.80E+04 

O § o o 

OO r - 1 

co o 00 

U-) ^ KT\ 

r- co 

j* 

't 't 

off f O 
w w w w w 

w, m vo o* o- 

r» ® in 

d od voi vd d 

1 .59E+05 
1.25E+05 
1.06E+05 

7.84E+04 

3.80E+04 

7.70E+04 

1.60E+05 

1.30E+05 

8.60E+04 

6.50E+04 

6.50E+04 

in in in 
O Q O 

+ T + 

WWW 
m o — 

o- r^i 

-5 1 

1 .08E+05 
1 .06E+05 
7.98E+04 

8.73E+04 

6.47E+04 

3.35E+05 

1.49E+05 

1.48E+05 

1 , 1 0E+05 

1.07B+05 

1.06E+05 

M5E+05 

inn 5 ^ 

??? f 

uimn u 
S38 8 

— « t— od od 

n in Tf 

?? ? 
W U UJ 
oo n n 
— 04 O 

W W ov 

B 

*?cs 

og? ° = 
~ ov 

|£5 Tf 3 w-i 

?f « ■ 

W tq ^ o S t w 

2^2 ° SS2 

7M «? — 

<N 

0^0 O 

CO 04 

cn m 

co ° in 
W-j 04 

X 

vl * 

>0 IO ^ ^ Tt 

??? f ? 

UJ W W U UJ 
© vo m m r- 
— o< — tt 

W — ’ 00 00 sd 

2.25E+05 

1.77E+05 

1.49E+05 

L10E+05 

5.40E+04 

1.09E+05 

2.25E+05 

1.60E+05 

1.06E+05 

8.00E+04 

8.00E+04 

in m, tj- 
C Q © 

+ T + 

WWW 
mm o 

04 04 m 

— : w d 

<N 

1 

0.0004804 

0.000661 

0.0003604 

0.000397 

0.0002945 

0.00153 

0.000675 

0.0006733 

0.0005 

0.000485 

0.0004825 

0.0005 

0.001 

0.0006395 

0.00049 

0.00049 

^ E 2 
§S 8 

d O o 

■ 

9.25E-06 

1.78E-04 

1.32E-05 

0 

0 

-1.01E-03 

0.00025 

1.34E-05 

0 

-0.00048 

2.50E-05 

0.001 

un 

o 

o§o o 

<N 

^ in 

o o 

W o w 
oo O 

P p 

W oo 

X 

I 

CO 

0.00097 

0.0015 

7.34E-04 

7.94E-04 

5.89E-04 

2.05E-03 

0.0016 

0.00136 

0.001 

0.00049 

9.90E-04 

0.002 

CO 

<n 22 O' g\ 

ml 

O O' O' 
m m 04 

88 8 
odd 

CS 

<T 

0.0099075 

0.0097 

0.007368 

0.008 

0.006 

0.01053 

0.00462 

0.004703 

0.0038 

0.00335 

0.003362 

0.0035 

gS| 1 

Sgl 8 

° P o d 
o 

m m 

o> in ^ 
— r- m 

o o m 

8 § 8 
o o 

_c 

A 

m 

OC vO o 

22% ° ° 
Oo« 

o 04 

-7.06E-03 

0.00176 

9.40E-05 

0 

-0.0033 

1.76E-04 

0.007 

f 

ogo o 
o> 

s s 

W o w 
o w o 
in — ; 

vd m 

X 

PJ 

?s 

-K.. m in vo <n 

S M - — f 

P o o © o 

° o b o' b 

Tt — w-i vo xj- o\ Tj- 

OOO O O b O 

ocv P P °. o 

00 (S oo oo 
- -h co oo 

§§s 8 

■ 

<= § 
<y < h 

e | 
’8 S 

a-c 

*r v 

A "V V 

o 1! 11 

© * g 

v 1 1 

A *U*U 

^ II II 

v 1 1 

w 

A r V'V 

2 ii ii 

OJ 2 g 

v 1 J 

A - ^ 

^ II II 

O x x 

° 1 J 


m 

<D 

| 

"E. 

<D 

I 

00 




























































Primary Orientation 

Case Number 

Secondary 
Orientation 
A deg 

Friction 

Coefficient 

V 

Max Gx 
(Tangential 
normal 
stress, ksi) 

0 (A=0, y^O) 

0 

0.0 

99.4 

0 

0 

0.3 I 

106.8 

0 

0 

0.7 1 

116.9 

0 

50 

0.0 

103.9 

0 

50 

0.3 

109.9 

0 

50 

0.7 

118.2 

5 (A=0, 7=7.5) 

0 

0.0 

105.7 

5 

0 

0.3 

113.1 

5 

0 

0.7 

123.3 

5 

50 

0.0 

108.6 

5 

50 

0.3 

114.4 

5 

50 

0.7 

122.6 

30 (A=5.74, 7*=- 13.86) 

0 

0.0 

84.3 (min) 

30 

0 

0.3 

89.9 (min) 

30 

0 

0.7 

97.7 (min) 

30 

50 

0.0 

94.1 

30 

50 

0.3 

98.8 

30 

50 

0.7 

105.2 


Table 4 Values of crystal orientation, friction coefficient, and max tangential normal 
stress Ox at the critical contact location, for the 18 FE model runs 































Nagaraj K. Arakere and Gregory Swanson 

Fretting Stresses in Single Crystal Superalloy Turbine Blade Attachments 



320X 



Fig. 1 A subsurface fretting fatigue crack emanating from a carbide in a turbine blade 
attachment (PWA1422) and propagating along octahedral (111) shear planes [4], 









Strain Amplitude (Uniaxial LCF Data) 


Power Law Curve Fit ( R A 2 = 0.469) : Ae = 0.0238 N 


• • 


1000 10000 100000 1000000 
Cycles to Failure 


Fig. 3 Strain range Vs. Cycles to Failure for LCF test data (PWA1493 at 1200F) 



Max Shear Stress Amplitude on Slip Planes 


Power Law Curve Fit (R A 2 = 0.674~) : Ax = 397,758 N 



Fig. 4 Shear Stress Amplitude [AT max ] Vs. Cycles to Failure 



Fig. 5 Secondary Crystallographic Orientation, (3, Vs Crack Depth for the 
SSME AHPFTP 1st Stage Turbine Blade [8, 23] 








Fig. 7 Convention for Defining Crystal Orientation in Turbine Blades [8] 







a total of 297 material orientations. 




Case Number (Primary Orientation) 


80 91 

Theta, Degrees (Secondary Axis Orientation) 

Fig. 9 Maximum Shear Stress Amplitude (Atmax. ksi) Contour Plot at the 

Blade Tip Critical Point 



Theta, Degees (SeocndaryOriertaticn) 


Fig. 10 Normalized HCF life (Contour Plot) at the blade tip Critical Point, as a function 

of primary and secondary orientation 


Crystal Orientation: Case 2, 0 Primary, 22.5 Secondary, Von Mises, Suction Side 


(j! 


iUiiiiisi j 



fuel pump, assembly with refined MODELS, CASE 2 


ANSYS 5.2 

FEB 28 1997 

12:19:44 

NODAL SOLUTION 

STEP=3 

SUB =1 

TIME=3 

SEQV (AVG) 

DMX =.035574 
SMN =426.469 
SMX =369035 
SMXB=513816 
I , -12500 

[ — i 0 

md 12500 
HI 25000 
37500 
HI 50000 
62500 
75000 
87500 
100000 


CHI 


Crystal Orientation: Case 2, 0 Primary, 22.5 Secondary, Von Mises, Pressure Side 



ANSYS 5.2 

FEB 28 1997 

12:44:01 

NODAL SOLUTION 

STEP=3 

SUB =1 

TIME=3 

SEQV (AVG) 
DMX =.035574 
SMN =426.469 
SMX =369035 
SMXB=513816 
, , -12500 

rH 0 
mm ^2500 
25000 
pm 37500 

s iss 

™ 87500 

1 1 100000 


FUEL PUMP, ASSEMBLY WITH REFINED MODELS, CASE 2 


Fig. 11 Representative stress plots for the single crystal blade attachment region 








1 



Fig. 12 HPFTP/AT first stage blade vonMises stress plot with local zoom in of the suction side upper 
contact region at the blade leading edge and the local coordinate system used for the contact results. 



Fig. 13 Fretting/galling induced crack in the contact region (suction side, trailing edge of blade) 
Several arrest marks are visible. Crystal orientation: A = -6.7°, y= 1 1.3°, p =4.2°. 





r 


, 3 



<o 

TD 

• 

C/3 


<D 


P 

C/3 

on 



C/3 




o 

O 


c/3 rn 

.5 ii 
60 ^ 
*C 

° « 
<D i 

E ii 

1- 

^ o 

bo ■- 
P cj 

ii 

“ 13 

C 4— > 

1 & 

2 U 

U ^ 

T3 c 
<U o 

O • - 

*2 ° 

.S o 

bO iu 
C bO 
P5 t 3 

13 ^ 

bD W) 

’bb .5 

a ~~ 


tJ 

£ 

Pu 

rr 




bi 

• 

ta 


Case Number (Primary Orientation) 


Contour Plot of Max Shear Stress 



Theta, Degrees (Secondary Orientation) 


Fig. 15 Contour plot of max shear stress amplitude, AT max (ksi), at the critical contact 
location, as a function of primary (Case number) and secondary (P or 0) 
crystallographic orientation. 



Case Number (Primary Orientation) 


Contour Plot of t *yt 2 

max 



Theta, Degrees (Secondary Orientation) 


Fig. 16 Contour plot of the parameter, t max *(Ay/2), at the critical contact location, as a 
function of primary (Case number) and secondary (P or 0) 
crystallographic orientation. 



Case Number (Primary Orientation) 


Contour Plot of Max normal * Normal Strain/2 



1 1 1 1 1 i i i i 

0 10 20 30 40 50 60 70 80 90 

Theta, Degrees (Secondary Orientation) 


Fig. 17 Contour plot of the parameter, G ma x*(A£/2), at the critical contact location, as a 
function of primary (Case number) and secondary ((3 or 0) 
crystallographic orientation. 




o 

o o o o o o 

mvoooooooooo 
NOtOinOLOOOOOQ 
\rin^-oiMinr^oLnoino 
in ■Hcn-Hivxr'dorv.inC'JiM 

CM CO I C*> I I I I rif'l'iJO'rl 

II II II II I II II I II II II I II 



Dashed line indicates tl 
boundary of the contac 
between the blade and 
disc attachments, at the 
critical region.