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Effect of variable normal force and frequency on fretting wear response of Ti–6Al–4V contact
Wear 271 (2011) 1535–1542
Contents lists available at ScienceDirect
Wear journal homepage: www.elsevier.com/locate/wear
Effect of variable normal force and frequency on fretting wear response of Ti–6Al–4V contact B. van Peteghem, S. Fouvry ∗ , J. Petit LTDS – Ecole Centrale de Lyon, 36 av Guy de Collongue – 69130 Ecully, France
a r t i c l e
i n f o
Article history: Received 14 October 2010 Received in revised form 5 January 2011 Accepted 6 January 2011
Keywords: TI–6AL–4V Wear Variable normal force
a b s t r a c t Fretting wear damage is one of the major problems in turbojet compressors. In conventional fretting test investigations, the normal force is usually imposed constant whereas contact displacement oscillates. The real blade-disk contact is in fact much more complex: both normal force and relative sliding vary during the cycle. Hence, a new fretting wear system has been developed where the relative sliding evolution can be controlled and the normal force varied, and which includes a contact opening stage during the loading cycle. The wear processes and wear kinetics under constant and variable normal conditions have been investigated. We demonstrate that variation in the normal force and the contact opening sequence significantly modify the wear process. Another result of the normal force variation is that the dissipated energy decreases. In addition to the complex influence of the variable loading spectrum, the influence of frequency is investigated. It is shown that the formation of a titanium nitride layer previously observed in a constant normal force contact, which is related to limited oxygen diffusion inside the contact, no longer takes place in a contact opening situation where the fretting wear damage process is mainly controlled by an oxidation process. © 2011 Published by Elsevier B.V.
1. Introduction
2. Experimental procedure
In turbojet compressors, one limitation to the lifetime of a blade is fretting wear into the blade-disk contact. Many studies have been made to understand and quantify the damage in this contact. The influences of various parameters have been studied, such as contact size [1] or surface treatments [2,3]. In those studies, the normal force is always kept constant. However, in the blade-disk contact, the normal force is not constant and varies with displacement. This particular loading cycle is dependent on the contact configuration, shown in Fig. 1. During engine start-up, the blade enters into contact with the disk socket by centrifugal force, inducing a simultaneous rise of normal force and displacement. When the engine stops, the blade falls to the bottom of the socket, and the contact opens. To reproduce a similar loading cycle during fretting tests, a new fretting system has been developed at the LTDS to control the normal force during a fretting cycle. The material studied is an alpha-beta titanium Ti–6Al–4V. A cylinder-plane contact has been chosen as a compromise between simplicity and representativeness.
2.1. Material and contact type
∗ Corresponding author. Tel.: +33 4 72 18 65 62; fax: +33 4 78 33 11 40. E-mail address: [email protected] (S. Fouvry). 0043-1648/$ – see front matter © 2011 Published by Elsevier B.V. doi:10.1016/j.wear.2011.01.060
The tests have been performed on an alpha–beta titanium Ti–6Al–4V which is very frequently used in aeronautics, especially in turbojets. The mechanical characteristics of this material are given in Table 1. This material has formerly been studied in [1] to determine its wear kinetics for various contact sizes. To complete this size effect investigation, we tested an 80 mm radius cylinderon-plane configuration. The contact length is 8 mm. The normal force has been adjusted to reach a 525 MPa maximum contact pressure. For the variable normal force tests, this pressure is reached at the maximum normal force (see Fig. 4 for details).
2.2. Testing system The machine used is an MTS axial–torsion testing machine equipped with two servo-hydraulic actuators. Both normal force and tangential displacement are driven separately. As shown in Fig. 2, the normal force is driven by a radial servo-actuator and the displacement is driven by an axial servo-actuator. The tangential displacement is measured with a LVDT, the normal force is measured with a 25 kN sensor. During the test, the tangential force and the tangential displacement are recorded. The tangential force is measured with a 250 kN
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Nomenclature Ed Q Q* P, FN ˛V ı ıg ı*
dissipated energy tangential force maximum tangential force normal force applied on the contact energy wear rate displacement sliding displacement maximum displacement
Fig. 3. Definition of loading variables extracted from a fretting cycle analysis (constant normal force conditions).
Fig. 1. Blade-disk contact loading cycle.
Table 1 Mechanical properties of TI–6AL–4V. Elastic modulus E (GPa) Poisson’s coefficient Yield stress (MPa) Vickers hardness (HV0.3 ) Density
119 0.29 970 360 4.4
sensor. These variables are used to draw fretting loops for constant and variable normal force conditions. To compare the influence of a variable normal force on the fretting wear process, both constant and variable normal loading sequences have been implemented. Fig. 4 details the conditions which were tested, including a representative evolution of the fretting cycle. The maximum normal force was fixed at 8523 N, inducing a linear normal loading of 1065 N/mm, thus imposing a maximum Hertzian pressure of 525 MPa. For the constant normal
Fig. 2. Fretting wear setup.
force test, sinusoidal displacement was imposed with an amplitude adjusted to ı* = ±120 m, thus achieving a relative sliding amplitude around ıg * = ±75 m, the reference sliding amplitude which was previously investigated in [1]. The loading sequence to reproduce the blade-disk contact is much more complex. However it can be reproduced by applying square-shaped evolutions of P and ı. At the beginning of the loading cycle, both normal force and displacement increase linearly up to a maximum normal force P = 8523 N and total displacement amplitude ımax . Then they stabilize at a constant value during a given period before decreasing to zero at a similar rate as during the loading period. Note that if displacement is stabilized at zero, the normal force passes under slight compression (i.e. negative period) to reproduce the contact opening sequence which characterizes the blade disk loading. The Q–ı fretting cycle obtained for the variable normal force is very close to that in reference one, which is imposed in common turbojet compressor contacts. By adjusting the applied displacement amplitude, a similar total sliding amplitude within the interface could be obtained, thus rationalising the comparison between constant and variable normal loading condition. Indeed, as illustrated in Fig. 3, the sliding amplitude ıg is different to the applied displacement due to the tangential compliance of the contact and test apparatus. The tests were therefore monitored in order to generate a constant total stroke ıgmax = 2ıg * = 150 m whatever the loading configuration. Note that this total stroke corresponds to the wear investigations developed in [1] and corresponds to a gross slip regime favouring wear. The test system compliance is quasi constant therefore it was shown that the previous condition is achieved by imposing a quasi constant displacement ımax = 2ı* = 240 m. The studied loading conditions are compiled in Table 2. As described above, the loading conditions used in this study try to represent the typical loading in the engine. In the real contact, the engine vibrations generate micro-displacements. These displacements are not considered in this study. The effects of these micro-displacements are studied especially in [1]. A variable displacement test [4] was carried out under constant normal loading conditions, first to identify the sliding transition between partial and gross slip and to characterize the friction response. The variation of the friction coefficient during the vari-
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Fig. 4. Loading and displacement variation during one cycle. Constant normal force (a) and variable normal force (b).
Table 2 Test series. Test series
Normal force (N)
ımax (m)
Frequency (Hz)
Number of cycles
(a) (b) (c)
8523 (constant) 8523 (constant) Variable from 0 to 8 523
240 240 240
5 0.11; 1; 3; 5 0.11
1000–15,000 1000 1000–5000
able displacement test is shown in Fig. 5. The transition between gross slip and partial slip is identified by the discontinuity in friction coefficient. Hence at the sliding transition the friction coefficient is around t = 0.9, but it drops at stab = 0.7 when an established gross slip condition exists. Before and after the tests, the samples were weighted to determine mass loss. After the test, the samples were cleaned in ethanol by ultrasound. Then, the wear volume was measured using a 3D profilometer. The samples were also been visually analysed. The representative scars have been analysed by X-ray diffraction and scanning electron microscopy.
Three Table 2.
test
series
were
performed
as
described
in
3. Results and discussion 3.1. Tribological behaviour During the tests, we observed an evolution of the Q–ı fretting cycle. For constant normal force tests (a, b), there is a tangential force peak at the end of the contact during the first cycles, as shown in Fig. 6a. This peak is due to the accumulation of wear debris at the
Table 3 Test conditions. Test #
Number of cycles
Frequency (Hz)
Condition
2.ı∗g = ıgmax
Ed (J)
Wear volume V (m3 )
1 2 3 4 5 6 7 8 9 10 11 12 13 14
10,000 n/a 5000 7500 1000 1000 1000 1000 1000 2500 15,000 5000 2500 5000
5 n/a 5 5 0,11 5 0,11 1 3 0,11 5 0,11 0,11 0,11
Constant n/a Constant Constant Variable Constant Constant Constant Constant Variable Constant Variable Constant Constant
176.8 n/a 134 141 150 70 155 107 137.6 76.8 150.4 103 138 150
17,200 n/a 7010 10,400 563 1130 1330 878 1410 968 24,500 2920 3390 6670
1.81E+09 n/a 8.88E+08 1.14E+09 369E+08 1.17E+08 3.01E+07 4.24E+08 2.92E+08 5.46E+08 2.18E+09 1.26E+09 l.98E+09 3.61E+09
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Fig. 5. Variation of friction coefficient (+) and sliding amplitude (♦) during variable amplitude test.
edge of the contact. After a certain number of cycles, this peak is reduced and disappears, as in Fig. 6b. The (c) tests were performed with a variable normal force loading cycle. An example of Q–ı fretting loop is given in Fig. 7. Unlike the constant normal force loop, this one does not significantly vary with the number of cycles. This particular shape of fretting loop is very similar to those calculated for a real blade-disk contact. The evolution of the friction coefficient (COF) was recorded for every test. The evolution during the cycles is given in Fig. 8. During the first cycles, the COF increases until it reaches a stabilised value. The evolution of the COF for the various loadings can be considered to be the same. The dissipated energy for each fretting cycle is calculated by the integration of the fretting loop area, as shown in Eq. (1). Then,
Fig. 8. Evolution of the coefficient of friction versus number of cycles: () constant ) normal force, 5 Hz, (––) constant normal force, 0.11 Hz, ( variable normal force, 0.11 Hz.
the accumulated dissipated energy is determined by summing each fretting cycle contribution over the whole test duration (2).
Ed =
Q (t) · ı(t)dt
Ed =
N
Ed (i)
(1)
(2)
i=1
3.2. Wear analysis First, a quantitative analysis of wear was carried out. Table 3 compiles the tested conditions and the corresponding wear results. A first sequence of tests was performed by imposing a similar low
Fig. 6. Fretting loops for constant normal force loading at 0.11 Hz. (a) 250 cycles, (b) 3000 cycles.
Fig. 7. Fretting loops for variable normal force loading. (a) 250 cycles, (b) 3000 cycles.
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Fig. 9. Wear kinetics for constant normal force (×) and variable normal force (): ı∗g = 68 m, ıgMAX = 136 m, f = 0.11 Hz.
frequency f = 0.11 Hz. Because similar relative sliding strokes are imposed, this infers that similar mean sliding speeds were generated in the interfaces. 3.3. Wear rate analysis Fig. 9a compares the wear evolution extension versus the fretting cycles. Linear evolutions are observed, suggesting that wear extends proportionally to the loading cycles. However the wear rates are very different, the constant normal force condition promotes a wear kinetics nearly three times higher than for variable normal force conditions. This tendency could be interpreted by considering the friction work which is dissipated in the interface. Indeed, by comparing the fretting loops, it can be concluded that the fretting cycle induced by variable normal force conditions is significantly less dissipative than that generated by an equivalent constant normal force. When plotting the evolution of the wear volume versus Ed , linear evolutions are usually observed, which justifies the following proportional relationship: V = ˛V ·
Ed
(3)
with ˛V the energy wear factor related to shape of the linear approximation. Plotting the wear volumes versus the accumulated dissipated energy (Fig. 9b), it can be observed that all the results are aligned along a single master curve. Hence a single energy wear coefficient can be extracted whatever the normal loading configuration. ˛V = 534, 900 m3 /J
(4)
This result is very important because it confirms the potential interest of the energy wear approach to quantify the wear rate under very complex sliding conditions. Secondly, it demonstrates that the conventional constant normal force tests are still pertinent to establish the energy wear rate of complex blade disk sliding assuming that equivalent sliding speeds are imposed. In order to evaluate the influence of frequency, and potentially the sliding speed, the tests performed at 0.11 Hz are compared with the results obtained under constant normal force at 5 Hz in Fig. 10. It is interesting to note that increasing of the loading frequency significantly reduces the wear rate. When the frequency is increased by a factor of 45, the energy wear rate is reduced by a factor of 5. Due to technical aspects (i.e. control of the loading sequences), it was not possible to investigate variable normal fretting tests at 5 Hz. Therefore we cannot know if a similar frequency energy wear rate reduction can be observed for a variable normal force sequence. Dedicated investigations are currently underway to clarify this aspect. However, an important conclusion of this investigation is that applying variable normal force conditions does not
seem to influence the energy wear rate, the loading frequency and, potentially, the sliding speed significantly affects the fretting wear kinetics. Another important conclusion extracted from this analysis concerns the representativeness of the conventional fretting wear test. Indeed, most fretting wear laws are extracted from a rather high frequency test, between 5 and 20 Hz. By contrast, the real gross slip frequency imposed on the blade disk contact is very low, equivalent to the flight frequency. Therefore this investigation, performed for an uncommonly low frequency (0.11 Hz), closer to flying conditions, underlines that most of the current fretting wear investigations tend to critically underestimate the real wear rate. In order to provide a representative and conservative energy wear rate, the test must be performed at very low frequency. Otherwise, using an accelerated test, dangerously optimistic wear rates should be extracted. Why low frequencies (i.e. slow sliding speeds) induces a highest wear rate is not clear. Indeed, it is commonly admitted that increasing the sliding speed, higher friction energy flow within the interface, resulting in high contact temperatures. These higher temperatures are supposed to accelerate the tribo-oxidation process and consequently the wear rate. Bearing this hypothesis in mind, a test performed at a higher frequency than the real case will systematically produce a pessimistic and conservative wear law. This investigation seams to show the opposite. In fact, the tribo-oxidation process implies temperature but also the time required for oxygen to react with titanium. Hence, the most plausible explanation related to this sharp increase of the energy wear rate with the frequency reduction is that, by increasing the time that oxygen can react with the native titanium metal, a thicker titanium oxide layer is formed on the surface. Consequently,
Fig. 10. Energy wear kinetics depending on normal force loading conditions and fretting sliding frequency. ı∗g = 68 m:() 5 Hz constant normal force, ()0.11 Hz variable normal force, (×) 0.11 Hz constant normal force.
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Fig. 11. Optical observations and wear profiles of fretting scars after 5000 cycles: (a) constant normal force (5 Hz, Ed = 6670 J),(b) (variable normal force (0.11 Hz,
Ed = 7010 J), (b) constant normal force (0.11 Hz,
Ed = 2920 J).)
the amount of material removed during each cycle significantly increases. One consequence is a significant increase of the friction energy efficiency to generate wear. Further investigations are now required for a better understanding of this phenomenon, including Fick modeling and coupling Quinn’s theory of tribo-oxidation processes [5]. 3.4. Analysis of the fretting scar structure In addition to the quantitative analysis of fretting kinetics, the structure and interface composition of fretting scars has been investigated. These analyses include optical observations, SEM observations, EDS chemical analysis and also DRX spectrum. Fig. 11 shows a significant difference between constant and normal force conditions. The fretting scar (a) (i.e. 5 Hz constant normal force) displays three distinct zones: • The centre of the scar is bright and white. • Around the centre, there is a bright orange zone. • On the edge, the scar is dark, between grey and black. The fretting scar (b) (i.e. 0.11 Hz constant normal force) displays quite a similar structure, but the boundaries between the different zones are less marked. The observation of fretting scar (c) (i.e. 0.11 Hz variable normal force) is very different. A homogenous structure characterized by dark oxides is observed over the whole interface. The wear profiles presented in Fig. 11 show differences which could be explained by the behaviour of the Tribological Transformed Structure (TTS). For the (a) scar, the TTS is present in the
Fig. 12. XRD spectrum of test planes before and after fretting tests (r) TI–6AL–4V plane before fretting test (reference) (a) TI–6AL–4V plane after constant normal force test (5 Hz) (b) TI–6AL–4V plane after constant normal force test (0.11 Hz) (c) TI–6AL–4V plane after variable normal force test (0.11 Hz).
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The XRD analysis was carried out with a K-␣ copper generator ˚ The measurements were made with a silicon – lithium at 1.54 A. diode, every 0.05◦ , over 1.5 s. The spectra obtained are presented in Fig. 12. The XRD spectrum shows small differences between structures before and after tests, and between tests:
Fig. 13. EDX spectrum for constant normal force fretting scar centre.
centre of the scar, as shown by the W shape. For the (b) scar, the TTS has been removed due to the lower frequency. For (c) scar, no TTS has been formed. Those optical observations have been completed with X-ray diffraction analysis (XRD), scanning electron microscopy (SEM) and energy dispersive spectrometry (EDS).
• The (r) spectrum shows the structure of TI–6AL–4V. We observe the peaks of the ␣ structure, and one peak of the  structure. The absence of some  structure peaks could be explained by the texturing of the material and by the poor ability of the  phase to diffract X-rays. • The (a) spectrum shows the same peaks as the (r) one, but there are three other peaks of TiOx Ny . This shows a structure modification due to the appearance of TiOx Ny . • The (b) spectrum shows no differences in comparison to the (r) one. • The (c) spectrum shows no major differences in comparison to the (r) one. We can only observe the loss of the  structure peak. The above results could be compared to those obtained by Mary et al. in [6]. The presence of TiOx Ny was demonstrated for this contact configuration. However, the percentage of oxygen or nitrogen could not be measured using XRD analysis. So we have decided to use an SEM – EDS analysis to quantify the presence of oxygen and nitrogen.
Fig. 14. SEM observation and EDS semi-quantitative analysis of fretting scars Pmax = 525 MPa, ıgMAX = 2.ı∗g = 136 m, constant normal force at 0.11 Hz (c) variable normal force at l0.11 Hz.
N = 5000 cycles (a) constant normal force at 5 Hz (b)
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The SEM was used with a 20 kV tension and the EDS analysis was made over 60 s. An example of an EDS spectrum is given in Fig. 13. The results are shown in Fig. 14. One important conclusion extracted from this analysis is the fact that the formation of titanium nitride layer previously observed by Mary et al. in [6] for high pressure conditions is again observed for constant normal force, whatever the studied frequency (0.11; 5 Hz). This result is important because it confirms the hypothesis developed by Mary et al. that this formation of titanium nitride, which is observed in large, closed contacts (i.e. high pressure conditions) is mainly controlled by oxygen depletion of the air within the interface. The plastic strain rate directly related to the frequency seems not to affect this process. Even more interesting is the analysis of the variable normal force fretting scar. The interface mainly consists of an oxide layer without any evidence of a TiOx Ny structure, as shown by XRD and EDS spectrum. This result definitely confirms the oxygen depletion hypothesis previously developed to explain the formation of titanium nitride which is activated in closed interfaces. Indeed, variable normal force tests induce contact opening at each cycle, so air reaches the fretted interface. This favours a generalized, homogenous oxidation over the whole interface and prevents any titanium nitride formation activation at the centre of the contact. Hence, the various structures of the interface and the composition of the fretting scars can be explained: • Constant normal force: ◦ Composite structure of the fretting scar. ◦ Titanium nitride formation at the centre of the contact due to a progressive oxygen depletion process: only nitrogen reaches the centre of the contact. • Variable normal force: ◦ Homogenous interface. ◦ Generalized oxidation over the whole interface. No evidence of titanium nitride formation. 4. Conclusion This study was carried out on an innovative fretting setup which allowed us to investigate new areas, especially the influence of variable normal force. The tests compared constant normal force and variable normal force tests. The results were investigated using a multidisciplinary approach, combining physical and chemical analyses. First, we demonstrated that energy wear rate is not dependent
of the loading cycle. Then, two major effects have been demonstrated: • The test frequency controls the wear rate: the lower the frequency, the higher the energy wear rate. • The normal force sequence controls the interface structure and chemical composition: ◦ A constant normal force, inducing a closed interface, promotes formation of titanium nitride due to fretting. ◦ A variable normal force, inducing an opening contact, promotes a generalized oxidation process. In addition to these results, it seems that the normal force loading sequence, and consequently the possible activation of the formation of titanium nitride, does not affect the energy wear rate, at least not in this macro-scale investigation. Deeper local wear investigations will be made to complete this aspect. Using this innovative fretting wear setup, the conditions of titanium nitride formation have been specified and the preferred oxidation method has been demonstrated. Moreover, more representativeness of the blade-disk interface can be investigated to identify representative wear laws. Acknowledgements The authors would like to especially thank Safran Group, Snecma, for financial support and Engineer Bernard Beaugiraud for his precious help for material analyses. References [1] S. Fouvry, C. Paulin, S. Deyber, Impact of contact size and complex gross–partial slip conditions on Ti–6Al–4V/Ti–6Al–4V fretting wear, Tribology International 42 (3) (2009) 461–474. [2] Y. Fu, Improvement in fretting wear and fatigue resistance of Ti–6Al–4V by application of several surface treatments and coatings, Surface and Coatings Technology 106 (2–3) (1998) 193–197. [3] H. Lee, S. Mall, Fretting behavior of shot peened Ti–6Al–4V under slip controlled mode, Wear 260 (6) (2006) 642–651. [4] S. Heredia, S. Fouvry, Introduction of a new sliding regime criterion to quantify partial, mixed and gross slip fretting regimes: correlation with wear and cracking processes, Wear 269 (7–8) (2010) 515–524. [5] T. Quinn, The role of oxide films in the friction and wear behaviour of metals, microscopic aspects of adhesion and lubrication, in: Proceedings of the 34th International Meeting of the Société de Chimie Physique, Elsevier, 1981, pp. 579–597. [6] C. Mary, T. Mogne, B. Beaugiraud, B. Vacher, J. Martin, S. Fouvry, Tribochemistry of a Ti alloy under fretting in air: evidence of titanium nitride formation, Tribology Letters 34 (3) (2009) 211–222.
Contents lists available at ScienceDirect
Wear journal homepage: www.elsevier.com/locate/wear
Effect of variable normal force and frequency on fretting wear response of Ti–6Al–4V contact B. van Peteghem, S. Fouvry ∗ , J. Petit LTDS – Ecole Centrale de Lyon, 36 av Guy de Collongue – 69130 Ecully, France
a r t i c l e
i n f o
Article history: Received 14 October 2010 Received in revised form 5 January 2011 Accepted 6 January 2011
Keywords: TI–6AL–4V Wear Variable normal force
a b s t r a c t Fretting wear damage is one of the major problems in turbojet compressors. In conventional fretting test investigations, the normal force is usually imposed constant whereas contact displacement oscillates. The real blade-disk contact is in fact much more complex: both normal force and relative sliding vary during the cycle. Hence, a new fretting wear system has been developed where the relative sliding evolution can be controlled and the normal force varied, and which includes a contact opening stage during the loading cycle. The wear processes and wear kinetics under constant and variable normal conditions have been investigated. We demonstrate that variation in the normal force and the contact opening sequence significantly modify the wear process. Another result of the normal force variation is that the dissipated energy decreases. In addition to the complex influence of the variable loading spectrum, the influence of frequency is investigated. It is shown that the formation of a titanium nitride layer previously observed in a constant normal force contact, which is related to limited oxygen diffusion inside the contact, no longer takes place in a contact opening situation where the fretting wear damage process is mainly controlled by an oxidation process. © 2011 Published by Elsevier B.V.
1. Introduction
2. Experimental procedure
In turbojet compressors, one limitation to the lifetime of a blade is fretting wear into the blade-disk contact. Many studies have been made to understand and quantify the damage in this contact. The influences of various parameters have been studied, such as contact size [1] or surface treatments [2,3]. In those studies, the normal force is always kept constant. However, in the blade-disk contact, the normal force is not constant and varies with displacement. This particular loading cycle is dependent on the contact configuration, shown in Fig. 1. During engine start-up, the blade enters into contact with the disk socket by centrifugal force, inducing a simultaneous rise of normal force and displacement. When the engine stops, the blade falls to the bottom of the socket, and the contact opens. To reproduce a similar loading cycle during fretting tests, a new fretting system has been developed at the LTDS to control the normal force during a fretting cycle. The material studied is an alpha-beta titanium Ti–6Al–4V. A cylinder-plane contact has been chosen as a compromise between simplicity and representativeness.
2.1. Material and contact type
∗ Corresponding author. Tel.: +33 4 72 18 65 62; fax: +33 4 78 33 11 40. E-mail address: [email protected] (S. Fouvry). 0043-1648/$ – see front matter © 2011 Published by Elsevier B.V. doi:10.1016/j.wear.2011.01.060
The tests have been performed on an alpha–beta titanium Ti–6Al–4V which is very frequently used in aeronautics, especially in turbojets. The mechanical characteristics of this material are given in Table 1. This material has formerly been studied in [1] to determine its wear kinetics for various contact sizes. To complete this size effect investigation, we tested an 80 mm radius cylinderon-plane configuration. The contact length is 8 mm. The normal force has been adjusted to reach a 525 MPa maximum contact pressure. For the variable normal force tests, this pressure is reached at the maximum normal force (see Fig. 4 for details).
2.2. Testing system The machine used is an MTS axial–torsion testing machine equipped with two servo-hydraulic actuators. Both normal force and tangential displacement are driven separately. As shown in Fig. 2, the normal force is driven by a radial servo-actuator and the displacement is driven by an axial servo-actuator. The tangential displacement is measured with a LVDT, the normal force is measured with a 25 kN sensor. During the test, the tangential force and the tangential displacement are recorded. The tangential force is measured with a 250 kN
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Nomenclature Ed Q Q* P, FN ˛V ı ıg ı*
dissipated energy tangential force maximum tangential force normal force applied on the contact energy wear rate displacement sliding displacement maximum displacement
Fig. 3. Definition of loading variables extracted from a fretting cycle analysis (constant normal force conditions).
Fig. 1. Blade-disk contact loading cycle.
Table 1 Mechanical properties of TI–6AL–4V. Elastic modulus E (GPa) Poisson’s coefficient Yield stress (MPa) Vickers hardness (HV0.3 ) Density
119 0.29 970 360 4.4
sensor. These variables are used to draw fretting loops for constant and variable normal force conditions. To compare the influence of a variable normal force on the fretting wear process, both constant and variable normal loading sequences have been implemented. Fig. 4 details the conditions which were tested, including a representative evolution of the fretting cycle. The maximum normal force was fixed at 8523 N, inducing a linear normal loading of 1065 N/mm, thus imposing a maximum Hertzian pressure of 525 MPa. For the constant normal
Fig. 2. Fretting wear setup.
force test, sinusoidal displacement was imposed with an amplitude adjusted to ı* = ±120 m, thus achieving a relative sliding amplitude around ıg * = ±75 m, the reference sliding amplitude which was previously investigated in [1]. The loading sequence to reproduce the blade-disk contact is much more complex. However it can be reproduced by applying square-shaped evolutions of P and ı. At the beginning of the loading cycle, both normal force and displacement increase linearly up to a maximum normal force P = 8523 N and total displacement amplitude ımax . Then they stabilize at a constant value during a given period before decreasing to zero at a similar rate as during the loading period. Note that if displacement is stabilized at zero, the normal force passes under slight compression (i.e. negative period) to reproduce the contact opening sequence which characterizes the blade disk loading. The Q–ı fretting cycle obtained for the variable normal force is very close to that in reference one, which is imposed in common turbojet compressor contacts. By adjusting the applied displacement amplitude, a similar total sliding amplitude within the interface could be obtained, thus rationalising the comparison between constant and variable normal loading condition. Indeed, as illustrated in Fig. 3, the sliding amplitude ıg is different to the applied displacement due to the tangential compliance of the contact and test apparatus. The tests were therefore monitored in order to generate a constant total stroke ıgmax = 2ıg * = 150 m whatever the loading configuration. Note that this total stroke corresponds to the wear investigations developed in [1] and corresponds to a gross slip regime favouring wear. The test system compliance is quasi constant therefore it was shown that the previous condition is achieved by imposing a quasi constant displacement ımax = 2ı* = 240 m. The studied loading conditions are compiled in Table 2. As described above, the loading conditions used in this study try to represent the typical loading in the engine. In the real contact, the engine vibrations generate micro-displacements. These displacements are not considered in this study. The effects of these micro-displacements are studied especially in [1]. A variable displacement test [4] was carried out under constant normal loading conditions, first to identify the sliding transition between partial and gross slip and to characterize the friction response. The variation of the friction coefficient during the vari-
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Fig. 4. Loading and displacement variation during one cycle. Constant normal force (a) and variable normal force (b).
Table 2 Test series. Test series
Normal force (N)
ımax (m)
Frequency (Hz)
Number of cycles
(a) (b) (c)
8523 (constant) 8523 (constant) Variable from 0 to 8 523
240 240 240
5 0.11; 1; 3; 5 0.11
1000–15,000 1000 1000–5000
able displacement test is shown in Fig. 5. The transition between gross slip and partial slip is identified by the discontinuity in friction coefficient. Hence at the sliding transition the friction coefficient is around t = 0.9, but it drops at stab = 0.7 when an established gross slip condition exists. Before and after the tests, the samples were weighted to determine mass loss. After the test, the samples were cleaned in ethanol by ultrasound. Then, the wear volume was measured using a 3D profilometer. The samples were also been visually analysed. The representative scars have been analysed by X-ray diffraction and scanning electron microscopy.
Three Table 2.
test
series
were
performed
as
described
in
3. Results and discussion 3.1. Tribological behaviour During the tests, we observed an evolution of the Q–ı fretting cycle. For constant normal force tests (a, b), there is a tangential force peak at the end of the contact during the first cycles, as shown in Fig. 6a. This peak is due to the accumulation of wear debris at the
Table 3 Test conditions. Test #
Number of cycles
Frequency (Hz)
Condition
2.ı∗g = ıgmax
Ed (J)
Wear volume V (m3 )
1 2 3 4 5 6 7 8 9 10 11 12 13 14
10,000 n/a 5000 7500 1000 1000 1000 1000 1000 2500 15,000 5000 2500 5000
5 n/a 5 5 0,11 5 0,11 1 3 0,11 5 0,11 0,11 0,11
Constant n/a Constant Constant Variable Constant Constant Constant Constant Variable Constant Variable Constant Constant
176.8 n/a 134 141 150 70 155 107 137.6 76.8 150.4 103 138 150
17,200 n/a 7010 10,400 563 1130 1330 878 1410 968 24,500 2920 3390 6670
1.81E+09 n/a 8.88E+08 1.14E+09 369E+08 1.17E+08 3.01E+07 4.24E+08 2.92E+08 5.46E+08 2.18E+09 1.26E+09 l.98E+09 3.61E+09
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Fig. 5. Variation of friction coefficient (+) and sliding amplitude (♦) during variable amplitude test.
edge of the contact. After a certain number of cycles, this peak is reduced and disappears, as in Fig. 6b. The (c) tests were performed with a variable normal force loading cycle. An example of Q–ı fretting loop is given in Fig. 7. Unlike the constant normal force loop, this one does not significantly vary with the number of cycles. This particular shape of fretting loop is very similar to those calculated for a real blade-disk contact. The evolution of the friction coefficient (COF) was recorded for every test. The evolution during the cycles is given in Fig. 8. During the first cycles, the COF increases until it reaches a stabilised value. The evolution of the COF for the various loadings can be considered to be the same. The dissipated energy for each fretting cycle is calculated by the integration of the fretting loop area, as shown in Eq. (1). Then,
Fig. 8. Evolution of the coefficient of friction versus number of cycles: () constant ) normal force, 5 Hz, (––) constant normal force, 0.11 Hz, ( variable normal force, 0.11 Hz.
the accumulated dissipated energy is determined by summing each fretting cycle contribution over the whole test duration (2).
Ed =
Q (t) · ı(t)dt
Ed =
N
Ed (i)
(1)
(2)
i=1
3.2. Wear analysis First, a quantitative analysis of wear was carried out. Table 3 compiles the tested conditions and the corresponding wear results. A first sequence of tests was performed by imposing a similar low
Fig. 6. Fretting loops for constant normal force loading at 0.11 Hz. (a) 250 cycles, (b) 3000 cycles.
Fig. 7. Fretting loops for variable normal force loading. (a) 250 cycles, (b) 3000 cycles.
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Fig. 9. Wear kinetics for constant normal force (×) and variable normal force (): ı∗g = 68 m, ıgMAX = 136 m, f = 0.11 Hz.
frequency f = 0.11 Hz. Because similar relative sliding strokes are imposed, this infers that similar mean sliding speeds were generated in the interfaces. 3.3. Wear rate analysis Fig. 9a compares the wear evolution extension versus the fretting cycles. Linear evolutions are observed, suggesting that wear extends proportionally to the loading cycles. However the wear rates are very different, the constant normal force condition promotes a wear kinetics nearly three times higher than for variable normal force conditions. This tendency could be interpreted by considering the friction work which is dissipated in the interface. Indeed, by comparing the fretting loops, it can be concluded that the fretting cycle induced by variable normal force conditions is significantly less dissipative than that generated by an equivalent constant normal force. When plotting the evolution of the wear volume versus Ed , linear evolutions are usually observed, which justifies the following proportional relationship: V = ˛V ·
Ed
(3)
with ˛V the energy wear factor related to shape of the linear approximation. Plotting the wear volumes versus the accumulated dissipated energy (Fig. 9b), it can be observed that all the results are aligned along a single master curve. Hence a single energy wear coefficient can be extracted whatever the normal loading configuration. ˛V = 534, 900 m3 /J
(4)
This result is very important because it confirms the potential interest of the energy wear approach to quantify the wear rate under very complex sliding conditions. Secondly, it demonstrates that the conventional constant normal force tests are still pertinent to establish the energy wear rate of complex blade disk sliding assuming that equivalent sliding speeds are imposed. In order to evaluate the influence of frequency, and potentially the sliding speed, the tests performed at 0.11 Hz are compared with the results obtained under constant normal force at 5 Hz in Fig. 10. It is interesting to note that increasing of the loading frequency significantly reduces the wear rate. When the frequency is increased by a factor of 45, the energy wear rate is reduced by a factor of 5. Due to technical aspects (i.e. control of the loading sequences), it was not possible to investigate variable normal fretting tests at 5 Hz. Therefore we cannot know if a similar frequency energy wear rate reduction can be observed for a variable normal force sequence. Dedicated investigations are currently underway to clarify this aspect. However, an important conclusion of this investigation is that applying variable normal force conditions does not
seem to influence the energy wear rate, the loading frequency and, potentially, the sliding speed significantly affects the fretting wear kinetics. Another important conclusion extracted from this analysis concerns the representativeness of the conventional fretting wear test. Indeed, most fretting wear laws are extracted from a rather high frequency test, between 5 and 20 Hz. By contrast, the real gross slip frequency imposed on the blade disk contact is very low, equivalent to the flight frequency. Therefore this investigation, performed for an uncommonly low frequency (0.11 Hz), closer to flying conditions, underlines that most of the current fretting wear investigations tend to critically underestimate the real wear rate. In order to provide a representative and conservative energy wear rate, the test must be performed at very low frequency. Otherwise, using an accelerated test, dangerously optimistic wear rates should be extracted. Why low frequencies (i.e. slow sliding speeds) induces a highest wear rate is not clear. Indeed, it is commonly admitted that increasing the sliding speed, higher friction energy flow within the interface, resulting in high contact temperatures. These higher temperatures are supposed to accelerate the tribo-oxidation process and consequently the wear rate. Bearing this hypothesis in mind, a test performed at a higher frequency than the real case will systematically produce a pessimistic and conservative wear law. This investigation seams to show the opposite. In fact, the tribo-oxidation process implies temperature but also the time required for oxygen to react with titanium. Hence, the most plausible explanation related to this sharp increase of the energy wear rate with the frequency reduction is that, by increasing the time that oxygen can react with the native titanium metal, a thicker titanium oxide layer is formed on the surface. Consequently,
Fig. 10. Energy wear kinetics depending on normal force loading conditions and fretting sliding frequency. ı∗g = 68 m:() 5 Hz constant normal force, ()0.11 Hz variable normal force, (×) 0.11 Hz constant normal force.
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Fig. 11. Optical observations and wear profiles of fretting scars after 5000 cycles: (a) constant normal force (5 Hz, Ed = 6670 J),(b) (variable normal force (0.11 Hz,
Ed = 7010 J), (b) constant normal force (0.11 Hz,
Ed = 2920 J).)
the amount of material removed during each cycle significantly increases. One consequence is a significant increase of the friction energy efficiency to generate wear. Further investigations are now required for a better understanding of this phenomenon, including Fick modeling and coupling Quinn’s theory of tribo-oxidation processes [5]. 3.4. Analysis of the fretting scar structure In addition to the quantitative analysis of fretting kinetics, the structure and interface composition of fretting scars has been investigated. These analyses include optical observations, SEM observations, EDS chemical analysis and also DRX spectrum. Fig. 11 shows a significant difference between constant and normal force conditions. The fretting scar (a) (i.e. 5 Hz constant normal force) displays three distinct zones: • The centre of the scar is bright and white. • Around the centre, there is a bright orange zone. • On the edge, the scar is dark, between grey and black. The fretting scar (b) (i.e. 0.11 Hz constant normal force) displays quite a similar structure, but the boundaries between the different zones are less marked. The observation of fretting scar (c) (i.e. 0.11 Hz variable normal force) is very different. A homogenous structure characterized by dark oxides is observed over the whole interface. The wear profiles presented in Fig. 11 show differences which could be explained by the behaviour of the Tribological Transformed Structure (TTS). For the (a) scar, the TTS is present in the
Fig. 12. XRD spectrum of test planes before and after fretting tests (r) TI–6AL–4V plane before fretting test (reference) (a) TI–6AL–4V plane after constant normal force test (5 Hz) (b) TI–6AL–4V plane after constant normal force test (0.11 Hz) (c) TI–6AL–4V plane after variable normal force test (0.11 Hz).
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The XRD analysis was carried out with a K-␣ copper generator ˚ The measurements were made with a silicon – lithium at 1.54 A. diode, every 0.05◦ , over 1.5 s. The spectra obtained are presented in Fig. 12. The XRD spectrum shows small differences between structures before and after tests, and between tests:
Fig. 13. EDX spectrum for constant normal force fretting scar centre.
centre of the scar, as shown by the W shape. For the (b) scar, the TTS has been removed due to the lower frequency. For (c) scar, no TTS has been formed. Those optical observations have been completed with X-ray diffraction analysis (XRD), scanning electron microscopy (SEM) and energy dispersive spectrometry (EDS).
• The (r) spectrum shows the structure of TI–6AL–4V. We observe the peaks of the ␣ structure, and one peak of the  structure. The absence of some  structure peaks could be explained by the texturing of the material and by the poor ability of the  phase to diffract X-rays. • The (a) spectrum shows the same peaks as the (r) one, but there are three other peaks of TiOx Ny . This shows a structure modification due to the appearance of TiOx Ny . • The (b) spectrum shows no differences in comparison to the (r) one. • The (c) spectrum shows no major differences in comparison to the (r) one. We can only observe the loss of the  structure peak. The above results could be compared to those obtained by Mary et al. in [6]. The presence of TiOx Ny was demonstrated for this contact configuration. However, the percentage of oxygen or nitrogen could not be measured using XRD analysis. So we have decided to use an SEM – EDS analysis to quantify the presence of oxygen and nitrogen.
Fig. 14. SEM observation and EDS semi-quantitative analysis of fretting scars Pmax = 525 MPa, ıgMAX = 2.ı∗g = 136 m, constant normal force at 0.11 Hz (c) variable normal force at l0.11 Hz.
N = 5000 cycles (a) constant normal force at 5 Hz (b)
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The SEM was used with a 20 kV tension and the EDS analysis was made over 60 s. An example of an EDS spectrum is given in Fig. 13. The results are shown in Fig. 14. One important conclusion extracted from this analysis is the fact that the formation of titanium nitride layer previously observed by Mary et al. in [6] for high pressure conditions is again observed for constant normal force, whatever the studied frequency (0.11; 5 Hz). This result is important because it confirms the hypothesis developed by Mary et al. that this formation of titanium nitride, which is observed in large, closed contacts (i.e. high pressure conditions) is mainly controlled by oxygen depletion of the air within the interface. The plastic strain rate directly related to the frequency seems not to affect this process. Even more interesting is the analysis of the variable normal force fretting scar. The interface mainly consists of an oxide layer without any evidence of a TiOx Ny structure, as shown by XRD and EDS spectrum. This result definitely confirms the oxygen depletion hypothesis previously developed to explain the formation of titanium nitride which is activated in closed interfaces. Indeed, variable normal force tests induce contact opening at each cycle, so air reaches the fretted interface. This favours a generalized, homogenous oxidation over the whole interface and prevents any titanium nitride formation activation at the centre of the contact. Hence, the various structures of the interface and the composition of the fretting scars can be explained: • Constant normal force: ◦ Composite structure of the fretting scar. ◦ Titanium nitride formation at the centre of the contact due to a progressive oxygen depletion process: only nitrogen reaches the centre of the contact. • Variable normal force: ◦ Homogenous interface. ◦ Generalized oxidation over the whole interface. No evidence of titanium nitride formation. 4. Conclusion This study was carried out on an innovative fretting setup which allowed us to investigate new areas, especially the influence of variable normal force. The tests compared constant normal force and variable normal force tests. The results were investigated using a multidisciplinary approach, combining physical and chemical analyses. First, we demonstrated that energy wear rate is not dependent
of the loading cycle. Then, two major effects have been demonstrated: • The test frequency controls the wear rate: the lower the frequency, the higher the energy wear rate. • The normal force sequence controls the interface structure and chemical composition: ◦ A constant normal force, inducing a closed interface, promotes formation of titanium nitride due to fretting. ◦ A variable normal force, inducing an opening contact, promotes a generalized oxidation process. In addition to these results, it seems that the normal force loading sequence, and consequently the possible activation of the formation of titanium nitride, does not affect the energy wear rate, at least not in this macro-scale investigation. Deeper local wear investigations will be made to complete this aspect. Using this innovative fretting wear setup, the conditions of titanium nitride formation have been specified and the preferred oxidation method has been demonstrated. Moreover, more representativeness of the blade-disk interface can be investigated to identify representative wear laws. Acknowledgements The authors would like to especially thank Safran Group, Snecma, for financial support and Engineer Bernard Beaugiraud for his precious help for material analyses. References [1] S. Fouvry, C. Paulin, S. Deyber, Impact of contact size and complex gross–partial slip conditions on Ti–6Al–4V/Ti–6Al–4V fretting wear, Tribology International 42 (3) (2009) 461–474. [2] Y. Fu, Improvement in fretting wear and fatigue resistance of Ti–6Al–4V by application of several surface treatments and coatings, Surface and Coatings Technology 106 (2–3) (1998) 193–197. [3] H. Lee, S. Mall, Fretting behavior of shot peened Ti–6Al–4V under slip controlled mode, Wear 260 (6) (2006) 642–651. [4] S. Heredia, S. Fouvry, Introduction of a new sliding regime criterion to quantify partial, mixed and gross slip fretting regimes: correlation with wear and cracking processes, Wear 269 (7–8) (2010) 515–524. [5] T. Quinn, The role of oxide films in the friction and wear behaviour of metals, microscopic aspects of adhesion and lubrication, in: Proceedings of the 34th International Meeting of the Société de Chimie Physique, Elsevier, 1981, pp. 579–597. [6] C. Mary, T. Mogne, B. Beaugiraud, B. Vacher, J. Martin, S. Fouvry, Tribochemistry of a Ti alloy under fretting in air: evidence of titanium nitride formation, Tribology Letters 34 (3) (2009) 211–222.