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Notched behaviour of a silicon carbide particulate reinforced aluminium alloy matrix composite
MATERIALS SCIENCE & ENGINEERING ELSEVIER
Materials Science and EngineeringA215 (1996) 67-72
A
Notched behaviour of a silicon carbide particulate reinforced aluminium alloy matrix composite Ze Wen Huang 1, Ian R. McColl*, Samuel J. Harris Department of Materials Engineering and Materials Design, Universityof Nottingham, Nottingham NG7 2RD, UK Received I9 January 1996; revised 30 January 1996; accepted 18 March 1996
Abstract
The notched fatigue behaviour of a fine particulate reinforced A1-Cu-Mg (2124) matrix composite is compared with that of an equivalent monolithic alloy (2024), both in the T4 condition. The difference in response of the materials for tensile and fatigue loading is pronounced. Under tensile loading the composite exhibits 'notch weakening' and an almost brittle fracture, while the monolithic alloy shows a degree of 'notch strengthening'. In contrast, under fatigue loading the composite exhibits a much smaller notch sensitivity than that of the monolithic alloy. At 10 7 cycles the fatigue notch sensitivity factor is reduced by ~ 6.5. The notched tensile behaviour of the composite is interpreted in terms of its reduced ductility and susceptibility to cavitation. Conventional notch theory is able to account for the notch sensitivity of the monolithic alloy in fatigue, but not the composite. Keywords: Silicon carbide; Aluminiumalloy; Notched behaviour
1. Introduction In this paper the influence of a V-notch on the tensile and fatigue behaviour of a particulate reinforced aluminium alloy matrix composite (AMC217) is compared with that for an equivalent monolithic alloy (2024). The composite, which was produced by a powder metallurgical route, offers appreciable improvements in tensile and fatigue behaviour over the equivalent monolithic alloy, although these are obtained at the expense of reduced ductility [1-3] and an inability to be coldformed or welded without serious degradation of the local microstructure.
2. Experimental 2.1. Spec#nens and materials The composite (AMC217) comprised a 2124 matrix reinforced with 17 vol % of 3 lam silicon carbide particles. It was supplied as 15 mm diameter bar by * Corresponding author. aNow at: InterdisciplinaryResearch Centre in Materials for High Performance Applications, University of Birmingham, Edgbaston, Birmingham B15 2TT (UK).
Aerospace Metal Composites, Farnborough, UK. Manufacture by a powder metallurgical route involved mechanical alloying, hot vacuum degassing, hot isostatic pressing and hot extrusion. The 2024 alloy was ingot cast and extruded to 25 mm thickness. Alloy compositions are shown in Table 1. Specimen blanks (15 mm diameter) were heat treated to the T4 condition prior to machining to minimise the subsequent affects of macro-stresses induced by this treatment. Blanks were prepared from the 2024 alloy, aligned parallel with the direction of extrusion. The composite and 2024 alloy were solution treated at 505 °C and 495 °C, respectively, for 1 h, quenched in cold water and room temperature aged for > 100 h. Plain tensile and fatigue specimens were turned, respectively, to 6.00 or 4.00 mm (minimum) gauge diameter using new single point cemented carbide tooling. Fatigue specimens were polished, parallel to the speciTable i Material specification MateriaI
Alloycomposition (mass%)
2124 2024
4.2 Cu, 1.5 Mg, 0.60 Mn, <0.1 Si, <0.I Fe 4.3 Cu, 1.5 Mg, 0.65 Mn, 0.11 Si, 0.2i Fe
68
Z.W. Huang et al. / Materials Science and Engineering A215 (I996) 67-72
men axis, down to 1200 grade silicon carbide. Surface finish was assessed by stylus profilometry parallel to the specimen axis. Notched specimens were similar to the equivalent plain specimens, but the turned diameter was increased by 1.00 mm so that the 60° V-notch, 0.50 mm deep with a root radius of 0.15 mm, resulted in the same minimum gauge diameter. The V-notch was cut with a freshly profiled single point cemented carbide tool.
800 700 6OO
6
~. 500
5
v 400 u) u) ~) 300
4
O3
2.2. Tensile and fatigue tests
3
rn o -~. O D -'-",,
2
2OO 100
An extensometer of 25 mm gauge length was used during the initial part of each tensile test. Fatigue testing was carried out at a frequency of 10 Hz under load control. The ratio (R) of minimum fatigue stress (Vm~) to maximum fatigue stress ((Ymax) was set to 0.1. Self aligning grips were used. All tests were carried out in a normal laboratory enviromment, temperature 1822°C, relative humidity 45-55%. Fracture surfaces were assessed by scanning electron microscopy at a beam voltage of 20 kV.
Plain Notched Composite 16
800 II [ ] 700
.
.
.
.
.
.
14.7 []
.,.-.-, 600 -
FL(Y1) 0.2%PS (3('1) UTS (YI)
[] %EL
(Y2)
3. I. Tensile behaviour The plain composite showed appreciable improvements in tensile properties ( ~ 50%) over the monolithic alloy, while maintaining a respectable elongation to failure of 5.9% (Fig. 1). The notched composite showed little elongation, 1.4% over a 25 mm gauge length, and a degree of notch weakening. In contrast, the monolithic alloy showed a degree of notch strengthening. Fracture surfaces for the plain 2024 specimens exhibited equiaxed dimples (1-2 btm) and a few localised cracks, whilst those for the notched specimens showed widespread secondary cracking. Both plain and notched composite specimens showed a bimodal population of dimples, the larger dimples (3-4 ~tm) being associated with the reinforcing particles and the smaller dimples ( g 1 btm) with the matrix. The notched composite specimens also showed some secondary cracking. 3.2. Fatigue behaviour Fig. 2 shows a typical surface profile obtained from a composite specimen. All fatigue specimens exhibited a similar surface profile with an arithmetic surface roughness < 0.20 btm (4 mm traverse, 0.8 mm cutoff). S - N data are shown in Fig. 3. At 107 cycles the composite showed a plain fatigue strength of 430 MPa, a 150 MPa increment over the 2024 alloy. At low cycles the improvement was less marked. The notched fatigue
12
10
500-
3. Results
,14
'~" 400 1,~ 300-
6
60 2 0 0 -
4
8
Ill o ~ El fll ~°
O D -'---
100 -
i.. 0
0-
Plain Notched Monolithic Alloy Fig. 1. Tensile properties o£ plain and notched specimens.
strength of the composite was reduced by ~ 110 MPa at both low and high cycles, while that of the 2024 alloy was particularly severely affected at high cycles. Striation patterns were readily discernable in the fatigue crack growth regions of both plain and notched 2024 specimens, except close to the notch in the notched specimens. Overload fracture regions showed extensive dimpling. The plain composite specimens gen erally showed fatigue crack initiation at a number of sites (typically three) around the specimen perimeter, with adjacent smoother fracture regions accounting for < 10% of the total fracture face. These smoother re-
l
I
lmm
Fig. 2. Surface profile of a composite specimen after polishing.
Z.W. Huang et al./ Materials Science and Engineering .4215 (1996) 67-72 500 q
,
Composite-plain Composite-notched
Ol ~ o)
'
i
200 i ¢1I
0
~
"'--%,,
u]
F
1~o5o
69
Monolithic alloy-plarn
y~" ,.......... Monolithic aHoy-notched
100ooo ~Oooooo l06oeooo Number of cycles to failure
Fig. 3. S - N curves for plain and notched monolithic alloy and composite specimens, R = 0.1.
gions, in which both reinforcing particles and matrix striation patterns (Fig. 4a) were occasionally discernable are interpreted as the fatigue crack growth regions. Fatigue initiation sites could not be identified in the notched specimens. However, an ~ 100 ~m wide track Fig. 5. Notch root crack in: (a) a composite specimen, 1 x 107 cycles, <~.... =360 MPa, and (b) a monolithic alloy specimen, 2.1 x 104 cycles, cy~aX= 290 MPa.
around the periphery of the fracture face was appreciably smoother (Fig. 4b) than the remaining area (Fig. 4c) suggesting that fatigue initiation occurred at multiple sites. The relatively flat featureless appearance of this narrow track, which again accounted for < 10% of the fracture surface, might suggest that a degree of crack closure had occurred. The overload region for both the plain and notched (Fig. 4c) specimens showed a bimodal population of dimples as for the tensile specimens. Fig. 5a and b, respectively, show cracks in the notch root of unbroken composite and 2024 fatigue specimens. In the case of the composite the two major cracks, or it may be one crack, are associated with a reinforcing particle and/or surface damage induced during machining of the notch. The crack in the monolithic alloy is relatively wide with appreciable deformation apparent.
4. Discussion
4. I. Tensile behaviour
Fig. 4. Composite fracture face: (a) fatigue region of a plain specimen, 9.4 x 106 cycles, O'max 420 MPa, and (b) fatigue region and (c) overload region of a notched specimen, 1.23 x 105 cycles, c~ma× = 365 MPa. =
Due to the capacity of the monolithic alloy for plastic deformation it is not surprising that the introduction of the circumferential V-notch resulted in a degree of notch strengthening (Fig. t). This behaviour is associated with the triaxial state of stress in the notched region and the constraint placed on its defor-
70
Z.IV. Huang et al. / Materials Science and Eng#wering A215 (1996) 67-72
Table 2 Experimentally derived values of K r and Q No of cycles to failure
10 6
107
2024 alloy
Qc/QA
Composite
Kr
Q~
Kt-
Qc
2,5 3.5
0.7I 1.19
1.38 1.38
0.18 0.18
mation by the adjacent un-notched material, which experiences a lower true stress. The triaxial state of stress tends to promote void fmrnation and restrict notched ductility. The higher deformation stresses associated with notch strengthening would account for the widespread cracking of the notched fracture surface. The influence of the notch on the behaviour of the composite is much more severe resulting in notch weakening and an almost brittle failure. This notch sensitivity under monotonic loading is generally in agreement with the results of Boothby et al. [4] for Charpy impact measurements on the same composite. They report an order of magnitude decrease in impact fracture energy due to the introduction of a V-notch with a tip radius of 0.25 mm. The lower ductility of the composite would limit its capacity to deform and blunt the notch. This would be further exacerbated by the composite's greater susceptibility to void formation. Whitehouse et al. [5] have shown that during tensile testing of a particulate reinforced composite, cavities formed in the matrix adjacent to the reinforcement, along the line of the applied stress. Also, oxide stringers, which were aligned in the direction of previous processing and the applied stress, both promoted and tended to stabilise the voids, retarding coalescence and failure. Although their observations were for an ahimina reinforced aluminium composite the behaviour of the composite investigated here, with its higher flow stress, would not be expected to be too dissimilar. In the triaxial stress field associated with the notch, oxide stringers in the composite would not be aligned in the stress direction and would be expected to be less effective at stabilising voids. Also, this composite was commercially produced by a process in which considerable efforts are made to control the oxide content. A high density of cavitation nuclei in the triaxial stress field would account for the very limited notched ductility.
4.2. Fatigue behaviour The effect of notches on monolithic alloys has received considerable attention in the literature, see for example Smith et al. [6,7], Yates et al. [8] and Miller [9]. The notched fatigue behaviour is related to the geometry of the notch, and to the plain fatigue strength or threshold stress intensity range of the material. Cracks
0,25 0,15
are assumed to initiate rapidly in the plastic zone associated with the notch, grow, probably at a diminishing rate as they propagate out of this zone, and then either cease propagation or continue to grow dependent on the plastic zone associated with the crack itself. The notched fatigue strength % is related to the plain fatigue strength Cyp by: O'n = Kfo"p
(t)
where Kf is the fatigue strength reduction factor. Kr is related effectively to the geometry of the notch by [6]: Kf= [1 + 7.69~pD--1°'5
(2)
where D is the notch depth and 9 the notch root radius. Kf is normally smaller than the geometrical stress concentration factor .~ of the notch, In the case of sharply notched specimens ~n is alternatively predicted by [6]:
~Ko (Tn= 2N/-~
(3)
where 2Xfo is the threshold stress intensity range of the material. Fatigue notch sensitivities are compared using the fatigue notch sensitivity factor Q [10]: O=
Kr-1 Kt-1
(4)
From Fig. 3 the reduction in fatigue strength at 107 cycles due to the introduction of the notch, for which Kt = 3.1, is much smaller for the composite ( ~ 26%) than it is for the monolithic alloy ( ~ 72%), measured over the net section. Values of Kr and Q calculated from the data in Fig. 3 using Eqs. (1) and (4) are shown in Table 2. The fatigue notch sensitivity of the composite is smaller than that of the monolithic alloy by a factor of 6.5 at 107 cycles. Shimokawa et al. [11] have reported previously on an extensive statistical study of the notched fatigue behaviour of 2024-T4 sheet (1 mm in thickness) tested in four point bending. Extrapolation of their results to 107 cycles, for Kt=3.1, suggests a reduction in fatigue strength of v 57% which is comparable to the value reported on here ( ~ 72%) for the monolithic alloy. Substitution of appropriate geometrical values into Eq. (2) indicates a reduction in fatigue strength of 74% which is in close agreement with the observed value.
Z.W. Huang et al. / Materials Science and Engineer#zg A215 (1996) 67-72
Although reliable values of threshold stress intensity range are not available for the composite, Eq. (3) can be used to calculate such values, albeit using an equation which assumes a sharp notch. The measured fatigue strengths at 10 7 cycles (Fig. 3) suggest AKo values of ~ 3.1 and 14.5 MPa m m, respectively, for the 2024 alloy and composite. The value for the monolithic alloy is similar to that reported for crack growth rate observations on material in the T3 condition [12]. The plain fatigue strength of the composite could be accounted for almost entirely by its higher elastic modulus, 103 GPa compared with 69 GPa for the monolithic alloy. So that, neglecting local stress concentrations and residual stresses, the fatigue stress range experienced by the matrix would be ~ 67% of that applied to the composite. This reduction in stress range is comparable with the ratio of plain fatigue strengths of the monolithic alloy and composite at 107 cycles, that is ~ 65%. It could then be argued that the influence of the fine distribution of reinforcing particles is neutral, other than in reducing the stress range experienced by the matrix. Certainly, micrographs of the fatigue crack growth regions only occasionally reveal the reinforcing particles, although the cracks in the notch root shown in Fig. 5a do appear to be associated with a particle. However, the fatigue crack growth regions of the plain (and notched) composite specimens are very limited compared with the monolithic alloy, for the same endurance, suggesting that a much greater proportion of the fatigue life was occupied with short crack development. Clearly the modulus argument cannot account for this. Neither can it or the notch strength reduction factor calculated using Eq. (2) account for the relative notch insensitivity of the composite under fatigue loading. The latter is not surprising since the equation for Kf does not incorporate any microstructural parameters. Nevertheless, the relative notch insensitivity of the composite under fatigue loading is supported by the work of Maruyama et al. [13] who have reported the fretting fatigue performance of a fine particulate reinforced composite, for which notch sensitivity is an issue, to be superior to that of the equivalent monolithic alloy. Clearly a factor or factors, other than notch geometry, plain fatigue strength or enhanced elastic modulus, are playing a role in the notched fatigue performance of the composite. Although thermally induced residual macro-stresses in the near surface layers may have had a role their influence would be expected to have been small. The 15 mm diameter specimen blanks were fully heat treated prior to machining down to a minimum gauge diameter of 4.00 mm, although final machining may itself have introduced macro-stresses or a degree of work hardening. Micro-stresses due to the mis-match between the coefficient of thelmal expansion of the reinforcing particles and the matrix would be expected
7I
to result in residual tensile stresses in the matrix, certainly in regions adjacent to the reinforcing particles. So it could be argued that for all the composite fatigue tests reported on here the maximum stress in the matrix adjacent to the particles was equal to the local (work hardened) yield stress. The effects of changes in applied fatigue loading then being restricted to the minimum and mean fatigue stresses. If this were the case it would apply to the plain as well as notched composite fatigue tests and would suggest early crack initiation followed by a period, up to 107 cycles, during which the crack was essentially dormant. Even discounting the affects of micro-stresses, the stresses at the notch root in the matrix of the composite would have been cycling up to the local yield value so that early crack initiation would have been expected. A degree of crack closure may have occurred and there are indications of this in Fig. 4b. If early crack initiation did occur then the relative notch insensitivity must be associated with a reduced initial crack growth rate which would be consistent with the higher threshold stress intensity range for the composite inferred from Eq. (3). From Smith et al. [7] the size of the notch plastic zone, neglecting the influence of the reinforcing particles, is 0.13(D9) °5 giving a value of 36 ktm which is appreciably smaller than the estimated width (100 gm) of the fatigue crack growth zone for the notched composite, operated at its notched fatigue strength. The fatigue crack growth rate may well diminish as the crack propagates out of the stress field associated with the notch and then microstructural barriers, the particles and high density of grain boundaries [14], impede or trap the fatigue crack.
5. Conclusions a. The plain tensile and fatigue performance of the composite shows appreciable improvements over the equivalent monolithic alloy, except for elongation to failure. b. For the particular notch configuration and heat treatment regime reported on here the addition of reinforcing particles results in a notch weakened material under monotonic loading, but a much reduced notch sensitivity under fatigue loading (R = 0.1). c. Conventional theories on the effects of notches on fatigue strength are able to predict the reduction in strength of the monolithic alloy but not the composite. Other material or geometrical factors related to the reinforcing particles are required.
Acknowledgements The work described in this paper was funded by the EPSRC (grant no. GR/G33899). The authors thank
72
Z.W. Huang eta/. / Materials Science and Eng#wering A215 (1996) 67-72
Aerospace Metal Composites for the supply o f materials and M r A Sleigtholme for technical assistance.
References [1] S.J. Harris, I.R. McColl, Z.W. Huang, A. WigNns and J. Sears, Pote~Ttial use of recta/matrix composite in high speed machinery, Metal Matrix Composite lIT, Exp/oiti~zg the hzvestment, Institute of Metals, London, 1991. [2] Z.W. Huang, I.R. McColl and S.J. Harris, in C.T. Sun and T.T. Luo (eds.), Proc. 2nd hzt. Conf. on Composite Materials and Structures, Peking Uni. Press, Peking, I992, 846-852. [3] S.J. Harris, I.R. McColl and Z.W. Huang, in R.A.L. Drew and H. Mostaghaci (eds.), Proc. bzt. Syrup. on Developments and Applications of Ceramzcs and New Metal Alloys, Quebec, Canadian Institute of Mining, Metallurgy and Petroleum, 1993, 117126.
[4] R.M. Boothby and C.A. Hippsley, Mat. Sci. Teeh., I0 6 (1994) 565-571. [5] A.F. Whitehouse and T.W. Clyne, Mat. Sei. Teeh,, I0 6 (1994) 468-474. [6] R.A. Smith and K.J. Miller, Int. J. Mech. Sei., 20 (1978) 201-206. [7] R.A. Smith and K.3. Miller, Int. J. Mech. Sci., I9 (1977) 11-22. [8] J.R. Yates and M.W. Brown, Fatigue Fract. Eng. Mat. Strut., I0 3 (1987) 187-201. [9] K.J. Miller, ?roe. Inst. Mech. Eng., 205 (1991) 1-14. [I0] R.E. Peterson, in G. Sines and J.L.William (eds.) Notch Sensitivity, in Metal Fatigue, McGraw-Hill, New York, 1959, 293-306. [I1] T. Shimokawa and H. Hamaguchi, J. Eng. Mat. Tech., 107 (1985) 214-220. [12] S E Stanzl, H R Mayer and E K Tschegg, Mat. Sci. Eng., A 147 1 (1991) 45-54. [13] N. Maruyama, M. Sumita and K. Nakazawa, J. dpn. h~st. Metals, 57 ii (1993) 1268-1274. [I4] Z.W. Huang, I.R. McColl and S.J. Harris, Influence of secondary processing on a fine particulate reinforced aluminium alloy matrix composite, submitted for publication.
Materials Science and EngineeringA215 (1996) 67-72
A
Notched behaviour of a silicon carbide particulate reinforced aluminium alloy matrix composite Ze Wen Huang 1, Ian R. McColl*, Samuel J. Harris Department of Materials Engineering and Materials Design, Universityof Nottingham, Nottingham NG7 2RD, UK Received I9 January 1996; revised 30 January 1996; accepted 18 March 1996
Abstract
The notched fatigue behaviour of a fine particulate reinforced A1-Cu-Mg (2124) matrix composite is compared with that of an equivalent monolithic alloy (2024), both in the T4 condition. The difference in response of the materials for tensile and fatigue loading is pronounced. Under tensile loading the composite exhibits 'notch weakening' and an almost brittle fracture, while the monolithic alloy shows a degree of 'notch strengthening'. In contrast, under fatigue loading the composite exhibits a much smaller notch sensitivity than that of the monolithic alloy. At 10 7 cycles the fatigue notch sensitivity factor is reduced by ~ 6.5. The notched tensile behaviour of the composite is interpreted in terms of its reduced ductility and susceptibility to cavitation. Conventional notch theory is able to account for the notch sensitivity of the monolithic alloy in fatigue, but not the composite. Keywords: Silicon carbide; Aluminiumalloy; Notched behaviour
1. Introduction In this paper the influence of a V-notch on the tensile and fatigue behaviour of a particulate reinforced aluminium alloy matrix composite (AMC217) is compared with that for an equivalent monolithic alloy (2024). The composite, which was produced by a powder metallurgical route, offers appreciable improvements in tensile and fatigue behaviour over the equivalent monolithic alloy, although these are obtained at the expense of reduced ductility [1-3] and an inability to be coldformed or welded without serious degradation of the local microstructure.
2. Experimental 2.1. Spec#nens and materials The composite (AMC217) comprised a 2124 matrix reinforced with 17 vol % of 3 lam silicon carbide particles. It was supplied as 15 mm diameter bar by * Corresponding author. aNow at: InterdisciplinaryResearch Centre in Materials for High Performance Applications, University of Birmingham, Edgbaston, Birmingham B15 2TT (UK).
Aerospace Metal Composites, Farnborough, UK. Manufacture by a powder metallurgical route involved mechanical alloying, hot vacuum degassing, hot isostatic pressing and hot extrusion. The 2024 alloy was ingot cast and extruded to 25 mm thickness. Alloy compositions are shown in Table 1. Specimen blanks (15 mm diameter) were heat treated to the T4 condition prior to machining to minimise the subsequent affects of macro-stresses induced by this treatment. Blanks were prepared from the 2024 alloy, aligned parallel with the direction of extrusion. The composite and 2024 alloy were solution treated at 505 °C and 495 °C, respectively, for 1 h, quenched in cold water and room temperature aged for > 100 h. Plain tensile and fatigue specimens were turned, respectively, to 6.00 or 4.00 mm (minimum) gauge diameter using new single point cemented carbide tooling. Fatigue specimens were polished, parallel to the speciTable i Material specification MateriaI
Alloycomposition (mass%)
2124 2024
4.2 Cu, 1.5 Mg, 0.60 Mn, <0.1 Si, <0.I Fe 4.3 Cu, 1.5 Mg, 0.65 Mn, 0.11 Si, 0.2i Fe
68
Z.W. Huang et al. / Materials Science and Engineering A215 (I996) 67-72
men axis, down to 1200 grade silicon carbide. Surface finish was assessed by stylus profilometry parallel to the specimen axis. Notched specimens were similar to the equivalent plain specimens, but the turned diameter was increased by 1.00 mm so that the 60° V-notch, 0.50 mm deep with a root radius of 0.15 mm, resulted in the same minimum gauge diameter. The V-notch was cut with a freshly profiled single point cemented carbide tool.
800 700 6OO
6
~. 500
5
v 400 u) u) ~) 300
4
O3
2.2. Tensile and fatigue tests
3
rn o -~. O D -'-",,
2
2OO 100
An extensometer of 25 mm gauge length was used during the initial part of each tensile test. Fatigue testing was carried out at a frequency of 10 Hz under load control. The ratio (R) of minimum fatigue stress (Vm~) to maximum fatigue stress ((Ymax) was set to 0.1. Self aligning grips were used. All tests were carried out in a normal laboratory enviromment, temperature 1822°C, relative humidity 45-55%. Fracture surfaces were assessed by scanning electron microscopy at a beam voltage of 20 kV.
Plain Notched Composite 16
800 II [ ] 700
.
.
.
.
.
.
14.7 []
.,.-.-, 600 -
FL(Y1) 0.2%PS (3('1) UTS (YI)
[] %EL
(Y2)
3. I. Tensile behaviour The plain composite showed appreciable improvements in tensile properties ( ~ 50%) over the monolithic alloy, while maintaining a respectable elongation to failure of 5.9% (Fig. 1). The notched composite showed little elongation, 1.4% over a 25 mm gauge length, and a degree of notch weakening. In contrast, the monolithic alloy showed a degree of notch strengthening. Fracture surfaces for the plain 2024 specimens exhibited equiaxed dimples (1-2 btm) and a few localised cracks, whilst those for the notched specimens showed widespread secondary cracking. Both plain and notched composite specimens showed a bimodal population of dimples, the larger dimples (3-4 ~tm) being associated with the reinforcing particles and the smaller dimples ( g 1 btm) with the matrix. The notched composite specimens also showed some secondary cracking. 3.2. Fatigue behaviour Fig. 2 shows a typical surface profile obtained from a composite specimen. All fatigue specimens exhibited a similar surface profile with an arithmetic surface roughness < 0.20 btm (4 mm traverse, 0.8 mm cutoff). S - N data are shown in Fig. 3. At 107 cycles the composite showed a plain fatigue strength of 430 MPa, a 150 MPa increment over the 2024 alloy. At low cycles the improvement was less marked. The notched fatigue
12
10
500-
3. Results
,14
'~" 400 1,~ 300-
6
60 2 0 0 -
4
8
Ill o ~ El fll ~°
O D -'---
100 -
i.. 0
0-
Plain Notched Monolithic Alloy Fig. 1. Tensile properties o£ plain and notched specimens.
strength of the composite was reduced by ~ 110 MPa at both low and high cycles, while that of the 2024 alloy was particularly severely affected at high cycles. Striation patterns were readily discernable in the fatigue crack growth regions of both plain and notched 2024 specimens, except close to the notch in the notched specimens. Overload fracture regions showed extensive dimpling. The plain composite specimens gen erally showed fatigue crack initiation at a number of sites (typically three) around the specimen perimeter, with adjacent smoother fracture regions accounting for < 10% of the total fracture face. These smoother re-
l
I
lmm
Fig. 2. Surface profile of a composite specimen after polishing.
Z.W. Huang et al./ Materials Science and Engineering .4215 (1996) 67-72 500 q
,
Composite-plain Composite-notched
Ol ~ o)
'
i
200 i ¢1I
0
~
"'--%,,
u]
F
1~o5o
69
Monolithic alloy-plarn
y~" ,.......... Monolithic aHoy-notched
100ooo ~Oooooo l06oeooo Number of cycles to failure
Fig. 3. S - N curves for plain and notched monolithic alloy and composite specimens, R = 0.1.
gions, in which both reinforcing particles and matrix striation patterns (Fig. 4a) were occasionally discernable are interpreted as the fatigue crack growth regions. Fatigue initiation sites could not be identified in the notched specimens. However, an ~ 100 ~m wide track Fig. 5. Notch root crack in: (a) a composite specimen, 1 x 107 cycles, <~.... =360 MPa, and (b) a monolithic alloy specimen, 2.1 x 104 cycles, cy~aX= 290 MPa.
around the periphery of the fracture face was appreciably smoother (Fig. 4b) than the remaining area (Fig. 4c) suggesting that fatigue initiation occurred at multiple sites. The relatively flat featureless appearance of this narrow track, which again accounted for < 10% of the fracture surface, might suggest that a degree of crack closure had occurred. The overload region for both the plain and notched (Fig. 4c) specimens showed a bimodal population of dimples as for the tensile specimens. Fig. 5a and b, respectively, show cracks in the notch root of unbroken composite and 2024 fatigue specimens. In the case of the composite the two major cracks, or it may be one crack, are associated with a reinforcing particle and/or surface damage induced during machining of the notch. The crack in the monolithic alloy is relatively wide with appreciable deformation apparent.
4. Discussion
4. I. Tensile behaviour
Fig. 4. Composite fracture face: (a) fatigue region of a plain specimen, 9.4 x 106 cycles, O'max 420 MPa, and (b) fatigue region and (c) overload region of a notched specimen, 1.23 x 105 cycles, c~ma× = 365 MPa. =
Due to the capacity of the monolithic alloy for plastic deformation it is not surprising that the introduction of the circumferential V-notch resulted in a degree of notch strengthening (Fig. t). This behaviour is associated with the triaxial state of stress in the notched region and the constraint placed on its defor-
70
Z.IV. Huang et al. / Materials Science and Eng#wering A215 (1996) 67-72
Table 2 Experimentally derived values of K r and Q No of cycles to failure
10 6
107
2024 alloy
Qc/QA
Composite
Kr
Q~
Kt-
Qc
2,5 3.5
0.7I 1.19
1.38 1.38
0.18 0.18
mation by the adjacent un-notched material, which experiences a lower true stress. The triaxial state of stress tends to promote void fmrnation and restrict notched ductility. The higher deformation stresses associated with notch strengthening would account for the widespread cracking of the notched fracture surface. The influence of the notch on the behaviour of the composite is much more severe resulting in notch weakening and an almost brittle failure. This notch sensitivity under monotonic loading is generally in agreement with the results of Boothby et al. [4] for Charpy impact measurements on the same composite. They report an order of magnitude decrease in impact fracture energy due to the introduction of a V-notch with a tip radius of 0.25 mm. The lower ductility of the composite would limit its capacity to deform and blunt the notch. This would be further exacerbated by the composite's greater susceptibility to void formation. Whitehouse et al. [5] have shown that during tensile testing of a particulate reinforced composite, cavities formed in the matrix adjacent to the reinforcement, along the line of the applied stress. Also, oxide stringers, which were aligned in the direction of previous processing and the applied stress, both promoted and tended to stabilise the voids, retarding coalescence and failure. Although their observations were for an ahimina reinforced aluminium composite the behaviour of the composite investigated here, with its higher flow stress, would not be expected to be too dissimilar. In the triaxial stress field associated with the notch, oxide stringers in the composite would not be aligned in the stress direction and would be expected to be less effective at stabilising voids. Also, this composite was commercially produced by a process in which considerable efforts are made to control the oxide content. A high density of cavitation nuclei in the triaxial stress field would account for the very limited notched ductility.
4.2. Fatigue behaviour The effect of notches on monolithic alloys has received considerable attention in the literature, see for example Smith et al. [6,7], Yates et al. [8] and Miller [9]. The notched fatigue behaviour is related to the geometry of the notch, and to the plain fatigue strength or threshold stress intensity range of the material. Cracks
0,25 0,15
are assumed to initiate rapidly in the plastic zone associated with the notch, grow, probably at a diminishing rate as they propagate out of this zone, and then either cease propagation or continue to grow dependent on the plastic zone associated with the crack itself. The notched fatigue strength % is related to the plain fatigue strength Cyp by: O'n = Kfo"p
(t)
where Kf is the fatigue strength reduction factor. Kr is related effectively to the geometry of the notch by [6]: Kf= [1 + 7.69~pD--1°'5
(2)
where D is the notch depth and 9 the notch root radius. Kf is normally smaller than the geometrical stress concentration factor .~ of the notch, In the case of sharply notched specimens ~n is alternatively predicted by [6]:
~Ko (Tn= 2N/-~
(3)
where 2Xfo is the threshold stress intensity range of the material. Fatigue notch sensitivities are compared using the fatigue notch sensitivity factor Q [10]: O=
Kr-1 Kt-1
(4)
From Fig. 3 the reduction in fatigue strength at 107 cycles due to the introduction of the notch, for which Kt = 3.1, is much smaller for the composite ( ~ 26%) than it is for the monolithic alloy ( ~ 72%), measured over the net section. Values of Kr and Q calculated from the data in Fig. 3 using Eqs. (1) and (4) are shown in Table 2. The fatigue notch sensitivity of the composite is smaller than that of the monolithic alloy by a factor of 6.5 at 107 cycles. Shimokawa et al. [11] have reported previously on an extensive statistical study of the notched fatigue behaviour of 2024-T4 sheet (1 mm in thickness) tested in four point bending. Extrapolation of their results to 107 cycles, for Kt=3.1, suggests a reduction in fatigue strength of v 57% which is comparable to the value reported on here ( ~ 72%) for the monolithic alloy. Substitution of appropriate geometrical values into Eq. (2) indicates a reduction in fatigue strength of 74% which is in close agreement with the observed value.
Z.W. Huang et al. / Materials Science and Engineer#zg A215 (1996) 67-72
Although reliable values of threshold stress intensity range are not available for the composite, Eq. (3) can be used to calculate such values, albeit using an equation which assumes a sharp notch. The measured fatigue strengths at 10 7 cycles (Fig. 3) suggest AKo values of ~ 3.1 and 14.5 MPa m m, respectively, for the 2024 alloy and composite. The value for the monolithic alloy is similar to that reported for crack growth rate observations on material in the T3 condition [12]. The plain fatigue strength of the composite could be accounted for almost entirely by its higher elastic modulus, 103 GPa compared with 69 GPa for the monolithic alloy. So that, neglecting local stress concentrations and residual stresses, the fatigue stress range experienced by the matrix would be ~ 67% of that applied to the composite. This reduction in stress range is comparable with the ratio of plain fatigue strengths of the monolithic alloy and composite at 107 cycles, that is ~ 65%. It could then be argued that the influence of the fine distribution of reinforcing particles is neutral, other than in reducing the stress range experienced by the matrix. Certainly, micrographs of the fatigue crack growth regions only occasionally reveal the reinforcing particles, although the cracks in the notch root shown in Fig. 5a do appear to be associated with a particle. However, the fatigue crack growth regions of the plain (and notched) composite specimens are very limited compared with the monolithic alloy, for the same endurance, suggesting that a much greater proportion of the fatigue life was occupied with short crack development. Clearly the modulus argument cannot account for this. Neither can it or the notch strength reduction factor calculated using Eq. (2) account for the relative notch insensitivity of the composite under fatigue loading. The latter is not surprising since the equation for Kf does not incorporate any microstructural parameters. Nevertheless, the relative notch insensitivity of the composite under fatigue loading is supported by the work of Maruyama et al. [13] who have reported the fretting fatigue performance of a fine particulate reinforced composite, for which notch sensitivity is an issue, to be superior to that of the equivalent monolithic alloy. Clearly a factor or factors, other than notch geometry, plain fatigue strength or enhanced elastic modulus, are playing a role in the notched fatigue performance of the composite. Although thermally induced residual macro-stresses in the near surface layers may have had a role their influence would be expected to have been small. The 15 mm diameter specimen blanks were fully heat treated prior to machining down to a minimum gauge diameter of 4.00 mm, although final machining may itself have introduced macro-stresses or a degree of work hardening. Micro-stresses due to the mis-match between the coefficient of thelmal expansion of the reinforcing particles and the matrix would be expected
7I
to result in residual tensile stresses in the matrix, certainly in regions adjacent to the reinforcing particles. So it could be argued that for all the composite fatigue tests reported on here the maximum stress in the matrix adjacent to the particles was equal to the local (work hardened) yield stress. The effects of changes in applied fatigue loading then being restricted to the minimum and mean fatigue stresses. If this were the case it would apply to the plain as well as notched composite fatigue tests and would suggest early crack initiation followed by a period, up to 107 cycles, during which the crack was essentially dormant. Even discounting the affects of micro-stresses, the stresses at the notch root in the matrix of the composite would have been cycling up to the local yield value so that early crack initiation would have been expected. A degree of crack closure may have occurred and there are indications of this in Fig. 4b. If early crack initiation did occur then the relative notch insensitivity must be associated with a reduced initial crack growth rate which would be consistent with the higher threshold stress intensity range for the composite inferred from Eq. (3). From Smith et al. [7] the size of the notch plastic zone, neglecting the influence of the reinforcing particles, is 0.13(D9) °5 giving a value of 36 ktm which is appreciably smaller than the estimated width (100 gm) of the fatigue crack growth zone for the notched composite, operated at its notched fatigue strength. The fatigue crack growth rate may well diminish as the crack propagates out of the stress field associated with the notch and then microstructural barriers, the particles and high density of grain boundaries [14], impede or trap the fatigue crack.
5. Conclusions a. The plain tensile and fatigue performance of the composite shows appreciable improvements over the equivalent monolithic alloy, except for elongation to failure. b. For the particular notch configuration and heat treatment regime reported on here the addition of reinforcing particles results in a notch weakened material under monotonic loading, but a much reduced notch sensitivity under fatigue loading (R = 0.1). c. Conventional theories on the effects of notches on fatigue strength are able to predict the reduction in strength of the monolithic alloy but not the composite. Other material or geometrical factors related to the reinforcing particles are required.
Acknowledgements The work described in this paper was funded by the EPSRC (grant no. GR/G33899). The authors thank
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Z.W. Huang eta/. / Materials Science and Eng#wering A215 (1996) 67-72
Aerospace Metal Composites for the supply o f materials and M r A Sleigtholme for technical assistance.
References [1] S.J. Harris, I.R. McColl, Z.W. Huang, A. WigNns and J. Sears, Pote~Ttial use of recta/matrix composite in high speed machinery, Metal Matrix Composite lIT, Exp/oiti~zg the hzvestment, Institute of Metals, London, 1991. [2] Z.W. Huang, I.R. McColl and S.J. Harris, in C.T. Sun and T.T. Luo (eds.), Proc. 2nd hzt. Conf. on Composite Materials and Structures, Peking Uni. Press, Peking, I992, 846-852. [3] S.J. Harris, I.R. McColl and Z.W. Huang, in R.A.L. Drew and H. Mostaghaci (eds.), Proc. bzt. Syrup. on Developments and Applications of Ceramzcs and New Metal Alloys, Quebec, Canadian Institute of Mining, Metallurgy and Petroleum, 1993, 117126.
[4] R.M. Boothby and C.A. Hippsley, Mat. Sci. Teeh., I0 6 (1994) 565-571. [5] A.F. Whitehouse and T.W. Clyne, Mat. Sei. Teeh,, I0 6 (1994) 468-474. [6] R.A. Smith and K.J. Miller, Int. J. Mech. Sei., 20 (1978) 201-206. [7] R.A. Smith and K.3. Miller, Int. J. Mech. Sci., I9 (1977) 11-22. [8] J.R. Yates and M.W. Brown, Fatigue Fract. Eng. Mat. Strut., I0 3 (1987) 187-201. [9] K.J. Miller, ?roe. Inst. Mech. Eng., 205 (1991) 1-14. [I0] R.E. Peterson, in G. Sines and J.L.William (eds.) Notch Sensitivity, in Metal Fatigue, McGraw-Hill, New York, 1959, 293-306. [I1] T. Shimokawa and H. Hamaguchi, J. Eng. Mat. Tech., 107 (1985) 214-220. [12] S E Stanzl, H R Mayer and E K Tschegg, Mat. Sci. Eng., A 147 1 (1991) 45-54. [13] N. Maruyama, M. Sumita and K. Nakazawa, J. dpn. h~st. Metals, 57 ii (1993) 1268-1274. [I4] Z.W. Huang, I.R. McColl and S.J. Harris, Influence of secondary processing on a fine particulate reinforced aluminium alloy matrix composite, submitted for publication.