- Email: [email protected]
High cycle fatigue behaviour of microsphere Al2O3–Al particulate metal matrix composites
Available online at www.sciencedirect.com
Composites: Part B 39 (2008) 1257–1269 www.elsevier.com/locate/compositesb
High cycle fatigue behaviour of microsphere Al2O3–Al particulate metal matrix composites B.G. Park a, A.G. Crosky b,*, A.K. Hellier c a
National Research Laboratory, School of Materials Science and Engineering, Pusan National University, San 30, Jangjeon-dong, Keumjung-gu, Busan 609-735, Republic of Korea b School of Materials Science and Engineering, The University of New South Wales, Sydney, NSW 2052, Australia c Institute of Materials Engineering, Australian Nuclear Science and Technology Organisation, Private Mail Bag 1, Menai, NSW 2234, Australia Received 23 January 2006; received in revised form 7 January 2008; accepted 12 January 2008 Available online 15 February 2008
Abstract The high-cycle stress-life (S–N) curve and fatigue crack growth threshold (DKth) behaviour of COMRAL-85TM, a 6061 aluminium– magnesium–silicon alloy reinforced with 20 vol.% Al2O3-based polycrystalline ceramic microspheres, and manufactured by a liquid metallurgy route, have been investigated for a stress ratio of R = 1 (fully reversed loading). Fatigue testing was conducted on both smooth round bar (S–N) specimens and notched round bar (fatigue threshold) specimens. Unreinforced Al 6061-T6 also processed by a liquid metallurgy route and six powder metallurgy processed composites with particle volume fractions ranging between 5% and 30% were also studied. S–N data revealed that the powder metallurgy processed composites generally gave longer fatigue lives than the matrix alloy, whereas COMRAL-85TM exhibited a reduced fatigue life. The fatigue threshold results were very similar for all the composites, being lower than for Al 6061-T6. Fatigue failure mechanisms were determined from examination of the fracture surfaces and the crack profiles. Ó 2008 Elsevier Ltd. All rights reserved. Keywords: A. Metal matrix composites (MMCs); Particle-reinforcement; B. Fatigue; D. Fractography
1. Introduction Aluminium metal matrix composites (MMCs) have been of interest as engineering materials because of their higher specific strength and stiffness, as well as superior wear resistance, compared to unreinforced aluminium alloys. Particulate metal matrix composites (PMMCs) are of special interest owing to the low cost of their raw materials and their ease of fabrication, making them suitable for applications requiring relatively high volume production. The improved properties of PMMCs result from the addition of hard ceramic particles, the size, shape, volume fraction, interfacial properties and distribution of which determine the mechanical behaviour of the composite. However,
*
Corresponding author. Tel.: +61 (0)2 9385 4424; fax: +61 (0)2 9385 5956. E-mail address: [email protected] (A.G. Crosky). 1359-8368/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesb.2008.01.006
PMMCs have worse ductility, toughness and low-cycle fatigue properties than unreinforced alloys, limiting their usefulness in practice. The present work investigates the high-cycle stress-life (S–N) and fatigue crack growth threshold (DKth) behaviour for stress ratio R = 1 (fully reversed loading) of an aluminium 6061/alumina-based microsphere PMMC as a function of particulate volume fraction. The use of microspheres as a reinforcement differentiates this work from other similar studies. 2. Experimental methods 2.1. Experimental materials The PMMC studied here [1,2] was COMRAL-85TM (Comalco microsphere reinforced aluminium) consisting of an aluminium alloy 6061 matrix reinforced with 20 vol.% MICRAL-20TM (microsphere alumina), a fine-grained,
1258
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Nomenclature b p K KI N Nf R
notch radius of fatigue specimen applied load stress intensity factor mode I stress intensity factor number of cycles; life number of cycles to failure radius of parallel section of fatigue specimen; stress ratio
polycrystalline ceramic comprising two principal phases, corundum (a-Al2O3) and mullite (3Al2O3 2SiO2) in the ratio 32:68 (wt%), with a size range of 1–50 lm (average size 20 lm and a Weibull modulus of 15) [3]. The grains of each phase were typically less than 0.5 lm in size. COMRAL-85TM was manufactured by a molten metal mixing method. Preheated microspheres were directly inserted into the molten aluminium and the melt was vigorously stirred by means of an impeller. The composite was then cast in a mould 125 mm diameter by 149 mm in size. The unreinforced alloy was melted and cast into an ingot of the same size. In addition, composites were produced using a conventional powder metallurgy route. Metal powder and microspheres were blended, compacted by cold isostatic pressing, and then sintered into a 125 mm diameter by 357 mm billet. This was done for six different volume fractions of the MICRAL-20TM reinforcement, ranging from 5% to 30% in intervals of 5% (designated PM 5–PM 30). The powder metallurgy processed billets and both of the liquid metallurgy processed ingots were extruded into 19 mm diameter rods. Prior to extrusion, the dies and billets (or ingots) were preheated to 480 °C and the material then extruded at a speed of 8–10 m/min. There was no significant surface damage on the extruded rods. Only a single 1000 mm length of extruded rod was available for each of the powder metallurgy processed composites, and this limited the size and number of specimens that could be used in the experimental work. The heat treatment cycle consisted of solution treatment at 530 °C for 90 min, quenching into cold water, pre-aging for 20 h at room temperature, then artificial aging at 175 °C (designated T6). The aging curves had a similar shape for the two liquid metallurgy processed materials, with the hardness being higher for the composite. In both cases, the peak hardness occurred for an aging time of 12 h. However, the curves were relatively flat over the time range from 4 h to 12 h, and it was decided therefore to use the manufacturer’s recommended aging time of 8 h for both these materials for aging to the T6 condition. The aging curves were similar for all six of the powder metallurgy processed composites, with the hardness increasing progressively with volume fraction. Peak hardness was achieved in 6 h and this time was therefore used as the
S Vf a DK DKth Dr
stress volume fraction b/R stress intensity factor range threshold stress intensity factor range; fatigue crack growth threshold stress range
T6 aging time for these materials. Table 1 contains the hardness and tensile properties of all the materials used in this study. Longitudinal sections of the powder metallurgy and liquid metallurgy processed composites containing 20 vol.% reinforcement are shown in Fig. 1. Both composites showed little evidence of particle clustering, but the liquid metallurgy processed composite had a less homogeneous distribution of particles than the powder metallurgy processed material, exhibiting distinct particle-free bands running parallel to the extrusion direction. The inter-particle spacing was measured in the liquid metallurgy processed composite, and gave similar values both along and across the extrusion direction, these being 51 lm and 48 lm, respectively. 2.2. Fatigue test specimens Two types of fatigue testing were performed: testing for constructing an S–N curve and fatigue threshold testing. Blanks for the fatigue test specimens were cut from the 19 mm diameter extruded rod with the major stress axis parallel to the extrusion direction as shown in Fig. 2. Reduced size specimens were used because of the limited amount of material available. The fatigue specimens for the stress-life curve were smooth and cylindrical in the gauge section, which measured 3 mm in diameter. The specimens for the fatigue threshold testing were cylindrical
Table 1 Hardness and tensile properties of all materials used in the study (loading axis parallel to the extrusion direction) Material
Hardness (Hv)
Young’s modulus (GPa)
Yield stress (MPa)
UTS (MPa)
Elongation (%)
Al 6061-T6 PM 5 PM 10 PM 15 PM 20 COMRAL-85TM PM 25 PM 30
111 118 123 128 134 119a 138 142
68 77 81 84 89 88 93 95
293 332 326 322 309 311 320 318
315 357 351 347 339 341 353 343
10.0 9.7 8.2 6.2 5.3 3.7 3.3 2.5
a
Aged for 8 h. Peak hardness was 123 Hv after 12 h.
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
1259
2.4. Fatigue threshold testing Fatigue threshold values were obtained using the increasing DK (stress intensity factor range) method on the same testing machine. The fatigue loading condition was identical to the S–N type test, that is, R = 1. Testing was started at a very low load level and then stopped if the specimen was not broken after two million cycles (the runout condition). The load level was then increased and the testing continued. This procedure was carried out for each specimen in turn until it fractured. The K value was calculated using Eq. (1) below taken from Ref. [5]: p pffiffiffiffiffiffi b pbfI ðaÞ; a ¼ 2 R pb pffiffiffiffiffiffiffiffiffiffiffi 1 1 3 1 þ a þ a2 0:363a3 þ 0:731a4 1a fI ðaÞ ¼ 2 2 8
KI ¼
ð1Þ
where KI is the (mode I) stress intensity factor, p is the applied load, R is the radius of the parallel section of the specimen, and b is the notch radius of the specimen. Four specimens were tested for each material. The threshold value was obtained from the average between the DK value of the specimen which was fractured at the lowest DK range and the largest un-fractured DK value of the specimen which fractured at the highest DK range. 2.5. Examination of fracture surface and crack profile
Fig. 1. Optical micrographs of longitudinal sections of: (a) PM 20 and (b) COMRAL-85TM, at low magnification. Particle-free bands are clearly visible in the latter.
and 5 mm in diameter, but were notched to give a 3 mm diameter at the root of the notch. A schematic diagram of the specimens is shown in Fig. 2. The surfaces of the specimens for the S–N testing were ground with 600 grit silicon carbide paper to remove all circumferential scratches and machining marks. 2.3. Stress-life (S–N) testing To construct an S–N curve, fully reversed tension–compression fatigue testing (R = 1) was performed at constant stress control using a dynamic Instron 5200 servohydraulic fatigue testing machine with a frequency of 50 Hz. The stress ratio R can be defined as the minimum load divided by the maximum load in a cycle. Fatigue tests were terminated at complete separation of the specimen or after 107 cycles were accumulated (the run-out condition). The tests were conducted in a controlled laboratory air environment (relative humidity of 55%, temperature of 25 °C) and in accordance with ASTM E466-82 [4].
In the tension–compression fatigue test, the fracture surface was damaged during the compression cycles, tending to erase the fractographic features. To obtain fracture surfaces suitable for fractographic examination, a limited number of tension–tension fatigue tests were performed with a stress ratio, R = 0.2. Horizontal tension–tension fatigue testing was also carried out on a Schenck midget pulser at 2800 cycles/min with notched specimens and these were used for examining the crack path. During testing, fatigue crack propagation was monitored using a linear variable differential transformer (LVDT). As soon as the crack had propagated about 1 mm, testing was terminated. The fracture surfaces of the tension–tension fatigue specimens were examined using a JEOL 840 scanning electron microscope (SEM) at low magnification to identify the fatigue and final fracture regions, and at higher magnification in the fatigue region to identify areas of crack initiation and early crack growth. Standard preparation and imaging procedures were used. The specimens were given a coating of carbon and examined using an accelerating voltage of 20 kV. The fatigue cracks produced by horizontal tension–tension fatigue testing were also examined using the SEM. The specimens were sectioned longitudinally and metallographically polished to a colloidal silica suspension finish. The number of particles along the crack path was counted manually on micrographs taken using the SEM and compared to that present along an arbitrary path in a polished section
1260
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Fig. 2. Schematic drawing of fatigue specimens for the S–N type test and fatigue threshold test.
350 AA 6061 TM COMRAL-85
300 Stress Range (MPa)
in the bulk material. Debonded particles were weighted by a factor of two when obtaining the total particle count, since only 50% of the debonded particles should be present on either of the two mating fracture surfaces. The number of particles along an arbitrary path was obtained by counting the particles along five randomly selected straight lines parallel to the fracture surface and then averaging the results.
250 200 150 100 50
3. Results 3.1. Stress-life (S–N) curves S–N curves for the unreinforced alloy Al 6061-T6 and the liquid metallurgy processed composite COMRAL85TM are shown in Fig. 3. Only a minimal amount of data was obtained because of the limited availability of test material. The results indicated that the fatigue strength of the liquid metallurgy processed composite was lower than that of the unreinforced alloy for any given number of
0 4 10
5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 3. Fatigue life versus cyclic stress range for COMRAL-85TM and Al 6061-T6 (R = 1).
cycles. The difference was however reduced at lower stress levels. The S–N curves for the powder metallurgy processed composites are shown in Figs. 4–6. Again, only limited data could be obtained because of the limited availability
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269 350
350
PM 5
PM 25 300
Stress Range (MPa)
Stress Range (MPa)
300 250 200 150 100 50
5
10
6
10 Cycles to Failure (Nf)
7
10
10
100
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure 5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
PM 30
300
Stress Range (MPa)
Stress Range (MPa)
300 250 200 150 100 50
5
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 4. Fatigue life versus cyclic stress range for PM 5 and PM 10 (R = 1).
350
PM 15
300 250 200 150 100
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
5
10
200 150 100
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
6
10
250
50
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
Stress Range (MPa)
150
350
PM 10
6
10 Cycles to Failure (Nf)
7
10
350
8
10
PM 20
300
Stress Range (MPa)
200
0 4 10
8
350
250 200 150 100 50
250
50
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
50
1261
0 4 10
5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 6. Fatigue life versus cyclic stress range for PM 25 and PM 30 (R = 1).
of material. In general, the fatigue data were reasonably consistent at the higher stress levels but showed a pronounced degree of scatter as the stress level decreased. When all the data were combined, as shown in Fig. 7, no systematic effect of particle volume fraction on the stresslife behaviour of the composites was apparent. However, the fatigue life was generally better for the composites than for the unreinforced alloy, except at the highest stresses, with the improvement in fatigue life becoming more pronounced as the stress level became lower. PM 10 and PM 25 exhibited lower fatigue strengths than the rest of the powder metallurgy processed composites. The fatigue life data are compared for the liquid metallurgy and powder metallurgy processed composites which contained 20% volume fraction of microspheres, in Fig. 8. The fatigue life of the liquid metallurgy processed composite was substantially shorter than that of the powder metallurgy processed composite and, contrary to the results obtained for the powder metallurgy processed composites, it was also shorter than for the unreinforced alloy. 3.2. Fatigue crack growth thresholds
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 5. Fatigue life versus cyclic stress range for PM 15 and PM 20 (R = 1).
The results of the fatigue crack growth threshold tests for the unreinforced alloy and the liquid metallurgy processed composite are shown in Figs. 9 and 10, respectively. The fatigue crack growth threshold of the liquid metallurgy processed composite was lower than that of the
1262
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269 350 COMRAL-85
4 PM 10
Specimen No.
Stress Range (MPa)
300
TM
250 PM 5 PM 30 PM 20 PM 15
200
PM 25
3
2
Threshold Value: 3.5 ± 0.5 MPa√m
1 Solid Symbol: Unbroken Open Symbol: Broken
150 AA 6061
0 0
100
10
4
10
5
6
10
10
7
2
4
8
10
6
8
10
ΔK (MPa√m)
Cycles to Failure (N f)
Fig. 10. Fatigue threshold for COMRAL-85TM (R = 1).
Fig. 7. Stress-life curves which contain all powder metallurgy processed composites and Al 6061-T6 (R = 1).
PM 5
350
Specimen No.
3
Stress Range (MPa)
300
250 COMRAL-85
TM
2
Threshold Value: 3.64 MPa√m
1 PM 20
200
150
Solid Symbol: Unbroken Open Symbol: Broken
AA 6061
0
2
4
6
8
10
ΔK (MPa√m) 100 4 10
10
5
6
10
10
7
8
10
Cycles to Failure (N f)
PM 10
Fig. 8. Comparison of S–N curves for COMRAL-85 (R = 1).
and PM 20
p
unreinforced alloy p (3.50 ± 0.50 MPa m compared with 5.60 ± 1.97 MPa m). There was significant scatter in the results for the unreinforced alloy while the data for the liquid metallurgy processed composite were relatively consistent. The fatigue crack growth thresholds of the powder metallurgy processed composites are shown in Figs. 11–13 and a graph of DKth versus particle volume fraction is given in Fig. 14. The threshold values of these composites were also substantially lower than that of the unreinforced alloy.
AA 6061
Specimen No.
4
3
2
4
Specimen No.
TM
3
Threshold Value: 3.24 ± 0.41 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 11. Fatigue thresholds for PM 5 and PM 10 (R = 1).
However, there was no significant difference between the different volume fractions. The average values of DKth for the powder metallurgy processed and liquid metallurgy processed composites with the same volume fraction of reinforcement were the same, although the latter showed more scatter.
Threshold Value: 5.60 ± 1.97 MPa√m
3.3. Fractography 1 Solid Symbol: Unbroken Open Symbol: Broken
0 0
2
4
6
8
10
12
ΔK (MPa√m)
Fig. 9. Fatigue threshold for Al 6061-T6 (R = 1).
14
Fig. 15 shows the fatigue fracture surfaces of a number of the powder metallurgy processed composites at low magnification. For the 5% volume fraction composite, the fatigue fracture area was clearly visible on the fracture surface but there was no distinct fatigue fracture area in the
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
1263
9
PM 15
8
4
ΔKth (MPa√m)
Specimen No.
7 3 Threshold Value: 2.83 MPa√m
2
6 5 4 3 2
1 Solid Symbol: Unbroken Open Symbol: Broken
1 0
0
1
2
3
4
5
6
7
8
9
10
0
5
10
15
20
25
30
Particle Volume Fraction (%)
ΔK (MPa√m)
Fig. 14. Fatigue threshold versus particle volume fraction (R = 1). PM 20
Specimen No.
4
3
Threshold Value: 3.49 ± 0.29 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 12. Fatigue thresholds for PM 15 and PM 20 (R = 1).
Fig. 16 shows the microstructural features across the fatigue fracture surface of the powder metallurgy processed composite reinforced with 20% volume fraction of microspheres. The surface shows different microstructural features as DK increases. At near-threshold and low DK ranges, the surface was relatively featureless and flat. Only a few particles were revealed in this area. As DK increased, surface roughness was increased and more particles were revealed on the fracture surface. For higher DK, the surfaces were much rougher and similar to those of a tensile overload fracture. No fatigue striations were detected, as was also the case for all of the other composites investigated in this study.
PM 25
3.4. Crack propagation
Specimen No.
4
3
Threshold Value: 3.49± 0.29 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 17 shows a crack profile for the liquid metallurgy processed composite. Broken and debonded particles could be seen along the crack path. However, when the number of particles encountered on the crack path was compared with the number encountered on an arbitrary path, the number of particles associated with the crack profile was found to be considerably less than that encountered on the arbitrary path, Fig. 18. As shown in the micrograph of the crack tip area at high magnification, Fig. 17b, microvoids can be seen at the particle/matrix interface.
Specimen No.
PM 30
4
4. Discussion
3
4.1. Stress-life (S–N) behaviour Threshold Value: 2.83 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 13. Fatigue thresholds for PM 25 and PM 30 (R = 1).
composites with microspheres.
the
higher
volume
fractions
of
It is generally accepted that the fatigue lives of PMMCs are longer than those of unreinforced alloys when stresscontrolled fatigue tests are performed [6–11]. The main contribution to the increased fatigue life in PMMCs is the higher elastic modulus. The fatigue process is strongly dependent on the strain amplitude, and the higher elastic modulus in PMMCs results in lower total strain, and especially total plastic strain, at any given stress level [12]. However, the improvement in fatigue life is not the same for all stress levels. As shown in Fig. 7, the increase in fatigue strength for the powder metallurgy processed composites
1264
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Fig. 15. Fatigue fracture surfaces at low magnification (Dr = 300 MPa, R = 0.2): (a) PM 5, (b) PM 20 and (c) PM 30.
Fig. 16. Fatigue fracture surface of PM 20 (Dr = 240 MPa, R = 0.2).
was not significant (or non-existent) at high stress levels but became increasingly more pronounced as the applied fluctuating stress decreased. This is similar to the behaviour observed by Vyletel et al. [7] in 15 vol.% TiC reinforced AA 2124. In fatigue tests under fully reversed loading (R = 1), fatigue is controlled by the plastic strain in the high-stress low-cycle fatigue regime (Nf < 104) while it is controlled by the elastic strain in the low-stress high-cycle regime (Nf > 104) [13]. Vyletel et al. [7] have suggested that, at high stress levels, the improvement in fatigue life expected in the composite as a result of the decreased plastic strain was offset by a decrease in fatigue life due to cyclic ductility, so that no net improvement occurs. As the stress level decreases, plastic behaviour becomes less important. The elastic behaviour now becomes increasingly dominant
and the fatigue life of the composite shows a progressive improvement over that of the unreinforced alloy since the elastic strain amplitude is less in the composite because of the higher elastic modulus. As discussed above, the elastic modulus is considered to have an important effect on fatigue life, at least in the highcycle low-stress regime. It is surprising therefore that, while the addition of particulate to the unreinforced alloy increased its fatigue life markedly, there was no further improvement with increasing volume fraction, even though a progressive increase in modulus occurred [14,15]. This finding is, however, consistent with the results obtained by Masuda and Tanaka [16] in their study of fatigue in A356 aluminium alloy reinforced with 10% and 20% SiC particles. They again found that the fatigue strength was
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
1265
Fig. 17. Fatigue crack profile for COMRAL-85TM: (a) high magnification of the rectangular area and (b) crack tip area.
higher in the composites than in the unreinforced alloy but observed no further improvement with increasing volume fraction. An absence of improvement in fatigue strength with increasing volume fraction has also been reported by Hall et al. [9] from their study of AA 2124 reinforced with 20% and 35% SiC particles. They found that the two composites also had similar tensile strengths and attributed the lack of improvement in fatigue life to the beneficial effect of increased modulus being offset by a degradation due to particle fracture. This would also explain the behaviour observed in the present study. Both static strength [14,15] and fatigue strength were not increased with increasing volume fraction of particulate. The former is considered to be due to particle fracture and this could also explain the lack of improvement in fatigue performance. A number of workers have attributed improvements in fatigue performance to increased static strength [16–19]. The findings of the present study show a relationship between fatigue performance and static strength [14,15] and are thus consistent with this view. However, it is unclear why the fatigue performances of PM 10 and PM 25 are inferior to the rest. The liquid metallurgy processed composite showed a shorter fatigue life than the powder metallurgy processed composite containing the same volume fraction of particulate. Moreover, the fatigue life of the liquid metallurgy processed composite was shorter even than that of the unreinforced alloy. This difference in fatigue life between the liquid metallurgy and powder metallurgy processed composites is not surprising in view of the different processing routes. Material manufactured by the powder metallurgy process is expected to have better fatigue strength
than that manufactured by the liquid metallurgy process because of the more refined grain size [20]. Indeed, Turnbull and Rios [21] have shown that in unreinforced AA 5754 the endurance stress corresponding to N = 107 load cycles was proportional to the inverse square root of the grain size. The grain size of the powder metallurgy processed composites used in the present study was definitely finer than that of the liquid metallurgy processed composite, and this difference in grain size is considered to be responsible, at least in part, for the difference in fatigue behaviour. Differences in grain size may also be responsible for some of the difference in fatigue life observed between the powder metallurgy processed composites and the unreinforced alloy. Another possible reason for the inferior fatigue performance of the liquid metallurgy processed composite is the more inhomogeneous distribution of the reinforcing particles (see Fig. 1). Yu et al. [22] found that the fatigue strength of AA 6061 reinforced with SiC particles was lower than that of the unreinforced alloy and attributed this to particle clustering. Similarly, the poor fatigue performance of the liquid metallurgy processed composite compared with that of the unreinforced alloy may be a result of the inhomogeneous particle distribution in the former. Ref. [23] contains a comprehensive summary of highcycle S–N data mostly with R = 0.1 (but also R = 1) for various aluminium alloys discontinuously reinforced with alumina or silicon carbide. The only study directly relevant to this work is by Lloyd [24] and contains stress amplitude versus cycles to failure curves at R = 1 for extruded unreinforced 6061-T6, and 6061-T6 reinforced with 15 vol.% Al2O3 particles. The S–N curves for
1266
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
6061-T6 and the composite agree moderately well with our data at low stress/high cycles to failure but the shapes of the curves differ, especially for high stress/low cycles to failure, where Lloyd’s curves diverge while ours cross.
increasing volume fraction is due to particle fracture. It is noted that some void formation was observed at the particle/matrix interface in front of growing fatigue cracks and some contribution from this effect may also have occurred.
4.2. Fatigue crack initiation
4.3. Fatigue crack growth threshold
Fatigue life consists of two components: crack initiation life and crack growth life. In smooth specimen tests, the fatigue crack initiation life can occupy a significant part of the total fatigue life. In un-notched specimens of high static strength aluminium alloys, crack initiation usually occupies between 75% and 95% of the total fatigue life [25]. The proportion of life consumed by crack initiation also increases with increasing static tensile strength. PMMCs have much higher tensile strengths than unreinforced alloys and their fatigue life would therefore be expected to be controlled by fatigue crack initiation. These conclusions are supported by the findings of several researchers [26]. In the case of unreinforced aluminium alloys, the main crack initiation mechanism is considered to be the formation of persistent slip bands (PSBs) [27]. In precipitation hardened aluminium alloys, fatigue life is controlled by the growth kinetics of PSBs in the low-cyclic load range. It is generally accepted that the material within the PSB lacks strengthening precipitates and is much softer than the surrounding matrix. However, PMMCs have more diverse crack initiation sites which are due to the particle additions. Srivatsan et al. [28] suggested that the particle/matrix interface is an important crack initiation site. During cyclic deformation the mismatch that exists between the hard and brittle reinforcing particles and the ductile matrix favours the concentration of stress at and near the particle/matrix interface, causing the matrix in the immediate vicinity to fail prematurely or the particle to separate from the matrix. On the other hand, Llorca and Poza [29] concluded that the dominant damage mechanism under cyclic deformation was reinforcement fracture, particles being broken by cracks perpendicular to the loading axis. Besides this kind of intrinsic crack initiation site, there may be extrinsic crack initiation sites such as particle clusters, casting defects (in ingot metallurgy processed composites), and pores. The propensity for particle fracture depends on the particle strength and size. Large and weak particles have a greater probability of fracturing than small and strong particles [30]. The MICRAL-20TM reinforcement particles used in this study have a broad size range [3] from 1 to about 50 lm with a Weibull modulus of 15 and the particles, being polycrystalline, appeared to be inherently weak. Substantial particle fracture appeared to occur below the yield strength of the composites [14,15] and it is most likely that particle fracture made a major contribution to fatigue crack initiation. This is consistent with the conclusion that the lack of improvement observed in fatigue life with
Previous studies of the effect of particulate addition on fatigue threshold have given mixed results with some authors reporting an increase in DKth [31] and others reporting a decrease [32], as was found here. This contradictory behaviour appears to be due to differences in the material and in the testing technique. There are some important differences between the method used here and that used by others. In the present work, the material was obtained as small diameter extruded bar and this precluded the use of compact tension specimens, which have been used in most other work where the material has been obtained as plate. In this case, round specimens were used instead. In addition, the tests were performed at R = 1 whereas most other workers have used a low value of R (generally R < 0.3). It is significant, however,pthat Crawford [33] obtained a threshold value of 3.6 MPa m for the 20 vol.% liquid metallurgy processed composite COMRAL-85TMpwhich is in excellent agreement with the value of 3.5 MPa m obtained in the present study. Moreover, this result was obtained for R = 0.1 using a compact tension test specimen. This may indicate that the threshold behaviour is very dependent on the specific composite. The absence of any systematic change in DKth with volume fraction is surprising in view of the generally held belief that volume fraction is one of the parameters which influence fatigue threshold [34]. The results are, however, consistent with the findings of Sugimura and Suresh [32] who reported an absence of any significant change in DKth with volume fraction in Al–3.5Cu composites reinforced with SiC particles. It is noted that, as in the present study, DKth of the composite was found to be lower than that of the unreinforced alloy. In the present work, no difference was found in DKth between the cast and powder metallurgy processed composites, yet again differences are commonly reported [35]. These differences have been attributed to microstructural differences resulting in enhanced roughness-induced crack closure. However, when fatigue testing is performed under tension–compression conditions, that is R = 1, as in the present study, the crack closure effect is negligible. Kumai et al. [35] have confirmed that when the crack closure effect is removed, the DKth of cast composites is the same as that of powder metallurgy processed composites as found in the present study. Ref. [23] contains a comprehensive summary of fatigue crack growth data mostly with R = 0.1 (but also R = 1) for various aluminium alloys discontinuously reinforced with alumina or silicon carbide. The only study directly relevant to this work is by Lloyd [24] and contains fatigue
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
crack growth rate curves at R = 1 for 6061-T6, and 6061T6 reinforced with 15 vol.% Al2O3 particles. The threshold stress intensity range DKth for 6061-T6 is about p p 4.3 MPa m in Lloyd’s paper versus 5.60 ± 1.97 MPa m in our work, while p that for the 15% Al2O3 composite p is about 6.1 MPa m in Lloyd’s work versus 2.83 MPa m in ours. This higher threshold in Lloyd’s work is (wrongly) attributed to crack closure effects. 4.4. Fatigue crack propagation
Cumulative Number of Particles
As shown in Fig. 16, only a few particles were revealed on the fracture surface at the near-threshold and low DK ranges. This indicates that in the early stage of crack growth, the fatigue crack tends to avoid the particles and propagate through the matrix. This was also apparent from the measurements made along the crack profile which indicated that the crack propagated through a smaller number of particles than would be encountered along an arbitrary line (refer to Fig. 18). However as DK increased, more particles became evident on the fracture surface, indicating that the crack was starting to seek out fractured and debonded particles in front of the growing crack tip (see Figs. 16 and 17). The lack of fatigue striations on the surface at the mid-range of DK suggested that the fatigue crack propagated in a predominantly brittle manner. These findings are consistent with the results obtained by Kumai et al. [36] in their study of short and long fatigue crack propagation in AA 6061 alloy reinforced with 15 vol.% of SiC particles. They concluded that at low stress levels, the fatigue crack tended to avoid SiC particles and ran through the matrix. As the stress level increased, the particles began to crack and the fatigue crack then linked up with the cracked particles. Fatigue crack propagation is strongly dependent on the interaction between particles and the advancing crack tip. As shown by Padkin et al. [37], when the modulus of second phase particles in front of the advancing fatigue crack is higher than that of the matrix alloy, as in the present study, the crack tends to avoid the particles. However as 28 26 24 22 20 18 16 14 12 10 8 6 4 2 0
Random Path Crack Path
0
0.2
0.4
0.6
0.8
1
1.2
1.4
1.6
1.8
Distance from Notch (mm)
Fig. 18. Numbers of particles encountered on a random (circle symbol) and crack (square symbol) path.
1267
the crack grows, DK increases, so that the plastic zone size in front of the crack tip becomes larger and progressively more particles become included in the plastic zone. This increases the propensity for particle fracture and debonding. Some evidence of microvoid formation at the particle/ matrix interface was seen in the present study. This is consistent with in situ SEM observations of fatigue crack propagation in the liquid metallurgy processed composite COMRAL-85TM made by Hadianfard et al. [38] which showed that fatigue crack growth was strongly associated with the particle/matrix interface. 5. Conclusions The fatigue strength of the powder metallurgy processed composites was generally higher than that of the unreinforced alloy, as found in other studies on PMMCs. The improved performance of the composites is attributed to their higher elastic moduli. The beneficial effect of particle addition on fatigue strength was more evident at lower stress levels. There was no significant change in fatigue strength with increasing volume fraction of particulate. The absence of any effect of the particle volume fraction on fatigue strength is considered to be due to the beneficial effect of the increase in modulus with volume fraction being offset by a degradation due to particle fracture. The fatigue strength of the liquid metallurgy processed composite was lower than those of the powder metallurgy processed composites and also less than that of the unreinforced alloy. The poorer performance of the liquid metallurgy processed composite compared with the powder metallurgy processed composites is attributed to its coarser grain size. The DKth values for the composites were lower than for the unreinforced alloy, in agreement with the results of some prior studies but contrary to the results of others. There was no systematic change in the DKth value with particle volume fraction as found previously in SiC particle reinforced Al–3.5Cu. The DKth values for the powder metallurgy and liquid metallurgy processed composites with the same volume fraction (20%) of reinforcement were similar. This is attributed to the microstructural differences between the two types of composite becoming insignificant for the R value of 1 used in the present study. Growing fatigue cracks were found to avoid the particles, particularly at lower DK values, whereas growing cracks produced during static fracture were attracted to the particles. This difference in behaviour appears to be due to a difference in the process zone size. While the process zone is small, as in the early stages of fatigue, the growing crack is deflected by the particles due to their high modulus. As the process zone becomes larger, it incorporates progressively more particles and the propensity for particle fracture and decohesion increases.
1268
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Acknowledgements This work was supported by Comalco Research Centre (15 Edgars Road, Thomastown, VIC 3074, Australia) with additional financial assistance through an Australian Research Council small grant. The assistance and advice provided by Drs. M.J. Couper and K. Xia of CRC are gratefully acknowledged. Mr. R.P. Harrison of the Australian Nuclear Science and Technology Organisation (ANSTO) is thanked for the loan of a Schenck horizontal midget pulser fatigue testing machine. References [1] Couper MJ, Xia K. Development of microsphere reinforced metal matrix composites. In: Hansen N, Juul Jensen D, Leffers T, Lilholt H, Lorentzen T, Pedersen AS, Pedersen OB, editors. Proceedings of the twelfth Risø international symposium on metallurgy and materials science: metal matrix composites – processing, microstructure and properties. Risø National Laboratory, Roskilde, Denmark, September; 1991. p. 291–8. [2] Xia K. Interfaces and interfacial reactions in Comral composites. In: Bandyopadhyay S, Crosky AG, editors. Proceedings of the second Australian forum on metal matrix composites. University of New South Wales, Sydney, Australia, December; 1991. p. 1–2. [3] Comalco data sheets. Comalco Research Centre, Thomastown, Victoria, Australia; 1990. [4] ASTM E466-82, Standard practice for conducting constant amplitude axial fatigue tests of metallic materials. Philadelphia: American Society for Testing and Materials; 1982. [5] Murakami Y. Stress intensity factors handbook, 1st ed. vol. 1. Oxford: Pergamon; 1987. p. 40. [6] Bonnen JJ, Allison JE, Jones JW. Fatigue behavior of a 2xxx series aluminum alloy reinforced with 15 vol pct SiCp. Metall Trans A 1991;22A(5):1007–19. [7] Vyletel GM, Allison JE, Van Aken DC. The influence of matrix microstructure and TiC reinforcement on the cyclic response and fatigue behavior of 2219 Al. Metall Trans A 1993;24A(11):2545–57. [8] Harris SJ, Dinsdale K, Gao Y, Noble B. Influence of heat treatment on the monotonic and fatigue properties of aluminium alloy composites. In: Andersen SI, Lilholt H, Pedersen OB, editors. Proceedings of the ninth Risø international symposium on metallurgy and materials science: mechanical and physical behaviour of metallic and ceramic composites. Risø National Laboratory, Roskilde, Denmark, September; 1988. p. 373–82. [9] Hall JN, Jones JW, Sachdev AK. Particle size, volume fraction and matrix strength effects on fatigue behavior and particle fracture in 2124 aluminum–SiCp composites. Mater Sci Eng A 1994;A183(1–2): 69–80. [10] Liu J, Wang DZ, Yao CK, Yu WC, Wang ZG. Effects of ageing treatment on fatigue strength in a SiCw/6061 Al composite. J Mater Sci Lett 1992;11(14):970–2. [11] Shang JK, Ritchie RO. Fatigue of discontinuously reinforced metal matrix composites. In: Everett RK, Arsenault RJ, editors. Metal matrix composites. San Diego: Academic Press; 1991. p. 255–85. [12] Allison JE, Jones JW. Fatigue behaviour of discontinuously reinforced metal matrix composites. In: Suresh S, Mortensen A, Needleman A, editors. Fundamentals of metal–matrix composites. Boston: Butterworth; 1993. p. 275 [Chapter 15]. [13] Barsom JM, Rolfe ST. Fracture and fatigue control in structures: applications of fracture mechanics. 2nd ed. Englewood Cliffs, New Jersey: Prentice-Hall; 1987. p. 239. [14] Park BG, Crosky AG, Hellier AK. Material characterisation and mechanical properties of Al2O3–Al metal matrix composites. J Mater Sci 2001;36(10):2417–26.
[15] Park BG. Fracture and fatigue of Al2O3 particulate reinforced metal matrix composites. PhD thesis. University of New South Wales, Sydney, Australia; 1997. [16] Masuda C, Tanaka Y. Tensile and fatigue properties of SiC whiskers and SiC particulate reinforced aluminium composites. In: Proceedings of the seventh international conference on fracture (ICF7): advances in fracture research, vol. 4. Houston, Texas, USA, March; 1989. p. 3089–96. [17] Hurd NJ. Fatigue performance of alumina reinforced metal matrix composites. Mater Sci Technol 1988;4(6):513–7. [18] Mura T, Qin S, Fan H. A theory of fatigue crack initiation. In: Teoh SH, Lee KH, editors. Proceedings of the joint FEFG/ICF international conference on fracture of engineering materials and structures. Singapore, August, 1991. London and New York: Elsevier Applied Science; 1991. p. 57–68. [19] Harris SJ, Dinsdale K, Gao Y, Noble B. Influence of heat treatment on the monotonic and fatigue properties of aluminium alloy composites. In: Andersen SI, Lilholt H, Pedersen OB, editors. Proceedings of the ninth Risø international symposium on metallurgy and materials science: mechanical and physical behaviour of metallic and ceramic composites. Risø National Laboratory, Roskilde, Denmark, September; 1988. p. 378. [20] Allison JE, Jones JW. Fatigue behaviour of discontinuously reinforced metal–matrix composites. In: Suresh S, Mortensen A, Needleman A, editors. Fundamentals of metal–matrix composites. Boston: Butterworth; 1993. p. 269 [Chapter 15]. [21] Turnbull A, de los Rios ER. The effect of grain size on fatigue crack growth in an aluminium magnesium alloy. Fatigue Fract Eng Mater Struct 1995;18(11):1355–66. [22] Yu S-Y, Ishii H, Tohgo K. Corrosion fatigue of SiC whisker or SiC particulate reinforced 6061 aluminum alloy. Fatigue Fract Eng Mater Struct 1994;17(5):571–8. [23] Lewandowski JJ. Fracture and fatigue of particulate MMCs. In: Kelly A, Zweben C, editors-in-chief. Comprehensive composite materials, vol. 3. Oxford: Pergamon; 2000. p. 151–87 [Section 3.07]. [24] Lloyd DJ. Particle reinforced aluminium and magnesium matrix composites. Int Mater Rev 1994;39(1):1–23. [25] Hahn CT, Simon R. A review of fatigue crack growth in high strength aluminum alloys and the relevant metallurgical factors. Eng Fract Mech 1973;5(3):523–40. [26] Wang DZ, Liu J, Yao CK, He ZH, Geng L. The initiation of fatigue microcracks near SiC whisker ends in a silicon carbide whisker reinforced 6061 aluminium composite. J Mater Sci Lett 1994;13(13):957–9. [27] Baxter WJ, McKinney TR. Growth of slip bands during fatigue of 6061-T6 aluminum. Metall Trans A 1988;19A(1):83–91. [28] Srivatsan TS, Auradkar R, Prakash A, Lavernia EJ. Influence of microstructure on cyclic stress response and cyclic fracture behavior of aluminium alloy metal–matrix composite. In: Teoh SH, Lee KH, editors. Proceedings of the joint FEFG/ICF international conference on fracture of engineering materials and structures. Singapore, August, 1991. London and New York: Elsevier Applied Science; 1991. p. 194–202. [29] Llorca J, Poza P. Influence of reinforcement fracture on the cyclic stress–strain curve of metal–matrix composites. Acta Metall Mater 1995;43(11):3959–69. [30] Curry DA, Knott JF. Effects of microstructure on cleavage fracture stress in steel. Metal Sci 1978;12(11):511–4. [31] Llorca J, Ruiz J, Healy JC, Elices M, Beevers CJ. Fatigue crack propagation in salt water, air and high vacuum in a spray-formed particulate-reinforced metal matrix composite. Mater Sci Eng A 1994;A185(1–2):1–15. [32] Sugimura Y, Suresh S. Effects of SiC content on fatigue crack growth in aluminum alloys reinforced with SiC particles. Metall Trans A 1992;23A(8):2231–42. [33] Crawford BR. Fatigue of an aluminium-6061 matrix composite. PhD thesis. University of Queensland, St. Lucia, Australia; 1996.
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269 [34] Davidson DL. The growth of fatigue cracks through particulate SiC reinforced aluminum alloys. Eng Fract Mech 1989;33(6):965–77. [35] Kumai S, Hu J, Higo Y, Nunomura S. Effects of dendrite cell size and particle distribution on the near-threshold fatigue crack growth behaviour of cast Al–SiCp composites. Acta Mater 1996;44(6): 2249–57. [36] Kumai S, King JE, Knott JF. Short and long fatigue crack growth in a SiC reinforced aluminium alloy. Fatigue Fract Eng Mater Struct 1990;13(5):511–24.
1269
[37] Padkin AJ, Brereton MF, Plumbridge WJ. Fatigue crack growth in two-phase alloys. Mater Sci Technol 1987;3(3):217–23. [38] Hadianfard MJ, Healy JC, Mai Y-W. The influence of heat treatment on mechanical and fracture properties of an alumina microsphere reinforced aluminium matrix composite. In: Chandra T, Dhingra AK, editors. Proceedings of advanced composites ’93. University of Wollongong, Australia, February; 1993. p. 1199–1206.
Composites: Part B 39 (2008) 1257–1269 www.elsevier.com/locate/compositesb
High cycle fatigue behaviour of microsphere Al2O3–Al particulate metal matrix composites B.G. Park a, A.G. Crosky b,*, A.K. Hellier c a
National Research Laboratory, School of Materials Science and Engineering, Pusan National University, San 30, Jangjeon-dong, Keumjung-gu, Busan 609-735, Republic of Korea b School of Materials Science and Engineering, The University of New South Wales, Sydney, NSW 2052, Australia c Institute of Materials Engineering, Australian Nuclear Science and Technology Organisation, Private Mail Bag 1, Menai, NSW 2234, Australia Received 23 January 2006; received in revised form 7 January 2008; accepted 12 January 2008 Available online 15 February 2008
Abstract The high-cycle stress-life (S–N) curve and fatigue crack growth threshold (DKth) behaviour of COMRAL-85TM, a 6061 aluminium– magnesium–silicon alloy reinforced with 20 vol.% Al2O3-based polycrystalline ceramic microspheres, and manufactured by a liquid metallurgy route, have been investigated for a stress ratio of R = 1 (fully reversed loading). Fatigue testing was conducted on both smooth round bar (S–N) specimens and notched round bar (fatigue threshold) specimens. Unreinforced Al 6061-T6 also processed by a liquid metallurgy route and six powder metallurgy processed composites with particle volume fractions ranging between 5% and 30% were also studied. S–N data revealed that the powder metallurgy processed composites generally gave longer fatigue lives than the matrix alloy, whereas COMRAL-85TM exhibited a reduced fatigue life. The fatigue threshold results were very similar for all the composites, being lower than for Al 6061-T6. Fatigue failure mechanisms were determined from examination of the fracture surfaces and the crack profiles. Ó 2008 Elsevier Ltd. All rights reserved. Keywords: A. Metal matrix composites (MMCs); Particle-reinforcement; B. Fatigue; D. Fractography
1. Introduction Aluminium metal matrix composites (MMCs) have been of interest as engineering materials because of their higher specific strength and stiffness, as well as superior wear resistance, compared to unreinforced aluminium alloys. Particulate metal matrix composites (PMMCs) are of special interest owing to the low cost of their raw materials and their ease of fabrication, making them suitable for applications requiring relatively high volume production. The improved properties of PMMCs result from the addition of hard ceramic particles, the size, shape, volume fraction, interfacial properties and distribution of which determine the mechanical behaviour of the composite. However,
*
Corresponding author. Tel.: +61 (0)2 9385 4424; fax: +61 (0)2 9385 5956. E-mail address: [email protected] (A.G. Crosky). 1359-8368/$ - see front matter Ó 2008 Elsevier Ltd. All rights reserved. doi:10.1016/j.compositesb.2008.01.006
PMMCs have worse ductility, toughness and low-cycle fatigue properties than unreinforced alloys, limiting their usefulness in practice. The present work investigates the high-cycle stress-life (S–N) and fatigue crack growth threshold (DKth) behaviour for stress ratio R = 1 (fully reversed loading) of an aluminium 6061/alumina-based microsphere PMMC as a function of particulate volume fraction. The use of microspheres as a reinforcement differentiates this work from other similar studies. 2. Experimental methods 2.1. Experimental materials The PMMC studied here [1,2] was COMRAL-85TM (Comalco microsphere reinforced aluminium) consisting of an aluminium alloy 6061 matrix reinforced with 20 vol.% MICRAL-20TM (microsphere alumina), a fine-grained,
1258
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Nomenclature b p K KI N Nf R
notch radius of fatigue specimen applied load stress intensity factor mode I stress intensity factor number of cycles; life number of cycles to failure radius of parallel section of fatigue specimen; stress ratio
polycrystalline ceramic comprising two principal phases, corundum (a-Al2O3) and mullite (3Al2O3 2SiO2) in the ratio 32:68 (wt%), with a size range of 1–50 lm (average size 20 lm and a Weibull modulus of 15) [3]. The grains of each phase were typically less than 0.5 lm in size. COMRAL-85TM was manufactured by a molten metal mixing method. Preheated microspheres were directly inserted into the molten aluminium and the melt was vigorously stirred by means of an impeller. The composite was then cast in a mould 125 mm diameter by 149 mm in size. The unreinforced alloy was melted and cast into an ingot of the same size. In addition, composites were produced using a conventional powder metallurgy route. Metal powder and microspheres were blended, compacted by cold isostatic pressing, and then sintered into a 125 mm diameter by 357 mm billet. This was done for six different volume fractions of the MICRAL-20TM reinforcement, ranging from 5% to 30% in intervals of 5% (designated PM 5–PM 30). The powder metallurgy processed billets and both of the liquid metallurgy processed ingots were extruded into 19 mm diameter rods. Prior to extrusion, the dies and billets (or ingots) were preheated to 480 °C and the material then extruded at a speed of 8–10 m/min. There was no significant surface damage on the extruded rods. Only a single 1000 mm length of extruded rod was available for each of the powder metallurgy processed composites, and this limited the size and number of specimens that could be used in the experimental work. The heat treatment cycle consisted of solution treatment at 530 °C for 90 min, quenching into cold water, pre-aging for 20 h at room temperature, then artificial aging at 175 °C (designated T6). The aging curves had a similar shape for the two liquid metallurgy processed materials, with the hardness being higher for the composite. In both cases, the peak hardness occurred for an aging time of 12 h. However, the curves were relatively flat over the time range from 4 h to 12 h, and it was decided therefore to use the manufacturer’s recommended aging time of 8 h for both these materials for aging to the T6 condition. The aging curves were similar for all six of the powder metallurgy processed composites, with the hardness increasing progressively with volume fraction. Peak hardness was achieved in 6 h and this time was therefore used as the
S Vf a DK DKth Dr
stress volume fraction b/R stress intensity factor range threshold stress intensity factor range; fatigue crack growth threshold stress range
T6 aging time for these materials. Table 1 contains the hardness and tensile properties of all the materials used in this study. Longitudinal sections of the powder metallurgy and liquid metallurgy processed composites containing 20 vol.% reinforcement are shown in Fig. 1. Both composites showed little evidence of particle clustering, but the liquid metallurgy processed composite had a less homogeneous distribution of particles than the powder metallurgy processed material, exhibiting distinct particle-free bands running parallel to the extrusion direction. The inter-particle spacing was measured in the liquid metallurgy processed composite, and gave similar values both along and across the extrusion direction, these being 51 lm and 48 lm, respectively. 2.2. Fatigue test specimens Two types of fatigue testing were performed: testing for constructing an S–N curve and fatigue threshold testing. Blanks for the fatigue test specimens were cut from the 19 mm diameter extruded rod with the major stress axis parallel to the extrusion direction as shown in Fig. 2. Reduced size specimens were used because of the limited amount of material available. The fatigue specimens for the stress-life curve were smooth and cylindrical in the gauge section, which measured 3 mm in diameter. The specimens for the fatigue threshold testing were cylindrical
Table 1 Hardness and tensile properties of all materials used in the study (loading axis parallel to the extrusion direction) Material
Hardness (Hv)
Young’s modulus (GPa)
Yield stress (MPa)
UTS (MPa)
Elongation (%)
Al 6061-T6 PM 5 PM 10 PM 15 PM 20 COMRAL-85TM PM 25 PM 30
111 118 123 128 134 119a 138 142
68 77 81 84 89 88 93 95
293 332 326 322 309 311 320 318
315 357 351 347 339 341 353 343
10.0 9.7 8.2 6.2 5.3 3.7 3.3 2.5
a
Aged for 8 h. Peak hardness was 123 Hv after 12 h.
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
1259
2.4. Fatigue threshold testing Fatigue threshold values were obtained using the increasing DK (stress intensity factor range) method on the same testing machine. The fatigue loading condition was identical to the S–N type test, that is, R = 1. Testing was started at a very low load level and then stopped if the specimen was not broken after two million cycles (the runout condition). The load level was then increased and the testing continued. This procedure was carried out for each specimen in turn until it fractured. The K value was calculated using Eq. (1) below taken from Ref. [5]: p pffiffiffiffiffiffi b pbfI ðaÞ; a ¼ 2 R pb pffiffiffiffiffiffiffiffiffiffiffi 1 1 3 1 þ a þ a2 0:363a3 þ 0:731a4 1a fI ðaÞ ¼ 2 2 8
KI ¼
ð1Þ
where KI is the (mode I) stress intensity factor, p is the applied load, R is the radius of the parallel section of the specimen, and b is the notch radius of the specimen. Four specimens were tested for each material. The threshold value was obtained from the average between the DK value of the specimen which was fractured at the lowest DK range and the largest un-fractured DK value of the specimen which fractured at the highest DK range. 2.5. Examination of fracture surface and crack profile
Fig. 1. Optical micrographs of longitudinal sections of: (a) PM 20 and (b) COMRAL-85TM, at low magnification. Particle-free bands are clearly visible in the latter.
and 5 mm in diameter, but were notched to give a 3 mm diameter at the root of the notch. A schematic diagram of the specimens is shown in Fig. 2. The surfaces of the specimens for the S–N testing were ground with 600 grit silicon carbide paper to remove all circumferential scratches and machining marks. 2.3. Stress-life (S–N) testing To construct an S–N curve, fully reversed tension–compression fatigue testing (R = 1) was performed at constant stress control using a dynamic Instron 5200 servohydraulic fatigue testing machine with a frequency of 50 Hz. The stress ratio R can be defined as the minimum load divided by the maximum load in a cycle. Fatigue tests were terminated at complete separation of the specimen or after 107 cycles were accumulated (the run-out condition). The tests were conducted in a controlled laboratory air environment (relative humidity of 55%, temperature of 25 °C) and in accordance with ASTM E466-82 [4].
In the tension–compression fatigue test, the fracture surface was damaged during the compression cycles, tending to erase the fractographic features. To obtain fracture surfaces suitable for fractographic examination, a limited number of tension–tension fatigue tests were performed with a stress ratio, R = 0.2. Horizontal tension–tension fatigue testing was also carried out on a Schenck midget pulser at 2800 cycles/min with notched specimens and these were used for examining the crack path. During testing, fatigue crack propagation was monitored using a linear variable differential transformer (LVDT). As soon as the crack had propagated about 1 mm, testing was terminated. The fracture surfaces of the tension–tension fatigue specimens were examined using a JEOL 840 scanning electron microscope (SEM) at low magnification to identify the fatigue and final fracture regions, and at higher magnification in the fatigue region to identify areas of crack initiation and early crack growth. Standard preparation and imaging procedures were used. The specimens were given a coating of carbon and examined using an accelerating voltage of 20 kV. The fatigue cracks produced by horizontal tension–tension fatigue testing were also examined using the SEM. The specimens were sectioned longitudinally and metallographically polished to a colloidal silica suspension finish. The number of particles along the crack path was counted manually on micrographs taken using the SEM and compared to that present along an arbitrary path in a polished section
1260
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Fig. 2. Schematic drawing of fatigue specimens for the S–N type test and fatigue threshold test.
350 AA 6061 TM COMRAL-85
300 Stress Range (MPa)
in the bulk material. Debonded particles were weighted by a factor of two when obtaining the total particle count, since only 50% of the debonded particles should be present on either of the two mating fracture surfaces. The number of particles along an arbitrary path was obtained by counting the particles along five randomly selected straight lines parallel to the fracture surface and then averaging the results.
250 200 150 100 50
3. Results 3.1. Stress-life (S–N) curves S–N curves for the unreinforced alloy Al 6061-T6 and the liquid metallurgy processed composite COMRAL85TM are shown in Fig. 3. Only a minimal amount of data was obtained because of the limited availability of test material. The results indicated that the fatigue strength of the liquid metallurgy processed composite was lower than that of the unreinforced alloy for any given number of
0 4 10
5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 3. Fatigue life versus cyclic stress range for COMRAL-85TM and Al 6061-T6 (R = 1).
cycles. The difference was however reduced at lower stress levels. The S–N curves for the powder metallurgy processed composites are shown in Figs. 4–6. Again, only limited data could be obtained because of the limited availability
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269 350
350
PM 5
PM 25 300
Stress Range (MPa)
Stress Range (MPa)
300 250 200 150 100 50
5
10
6
10 Cycles to Failure (Nf)
7
10
10
100
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure 5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
PM 30
300
Stress Range (MPa)
Stress Range (MPa)
300 250 200 150 100 50
5
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 4. Fatigue life versus cyclic stress range for PM 5 and PM 10 (R = 1).
350
PM 15
300 250 200 150 100
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
5
10
200 150 100
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
6
10
250
50
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
Stress Range (MPa)
150
350
PM 10
6
10 Cycles to Failure (Nf)
7
10
350
8
10
PM 20
300
Stress Range (MPa)
200
0 4 10
8
350
250 200 150 100 50
250
50
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
50
1261
0 4 10
5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 6. Fatigue life versus cyclic stress range for PM 25 and PM 30 (R = 1).
of material. In general, the fatigue data were reasonably consistent at the higher stress levels but showed a pronounced degree of scatter as the stress level decreased. When all the data were combined, as shown in Fig. 7, no systematic effect of particle volume fraction on the stresslife behaviour of the composites was apparent. However, the fatigue life was generally better for the composites than for the unreinforced alloy, except at the highest stresses, with the improvement in fatigue life becoming more pronounced as the stress level became lower. PM 10 and PM 25 exhibited lower fatigue strengths than the rest of the powder metallurgy processed composites. The fatigue life data are compared for the liquid metallurgy and powder metallurgy processed composites which contained 20% volume fraction of microspheres, in Fig. 8. The fatigue life of the liquid metallurgy processed composite was substantially shorter than that of the powder metallurgy processed composite and, contrary to the results obtained for the powder metallurgy processed composites, it was also shorter than for the unreinforced alloy. 3.2. Fatigue crack growth thresholds
Solid Symbol: Terminated with failure Open Symbol: Terminated without failure
0 4 10
5
10
6
10 Cycles to Failure (Nf)
7
10
8
10
Fig. 5. Fatigue life versus cyclic stress range for PM 15 and PM 20 (R = 1).
The results of the fatigue crack growth threshold tests for the unreinforced alloy and the liquid metallurgy processed composite are shown in Figs. 9 and 10, respectively. The fatigue crack growth threshold of the liquid metallurgy processed composite was lower than that of the
1262
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269 350 COMRAL-85
4 PM 10
Specimen No.
Stress Range (MPa)
300
TM
250 PM 5 PM 30 PM 20 PM 15
200
PM 25
3
2
Threshold Value: 3.5 ± 0.5 MPa√m
1 Solid Symbol: Unbroken Open Symbol: Broken
150 AA 6061
0 0
100
10
4
10
5
6
10
10
7
2
4
8
10
6
8
10
ΔK (MPa√m)
Cycles to Failure (N f)
Fig. 10. Fatigue threshold for COMRAL-85TM (R = 1).
Fig. 7. Stress-life curves which contain all powder metallurgy processed composites and Al 6061-T6 (R = 1).
PM 5
350
Specimen No.
3
Stress Range (MPa)
300
250 COMRAL-85
TM
2
Threshold Value: 3.64 MPa√m
1 PM 20
200
150
Solid Symbol: Unbroken Open Symbol: Broken
AA 6061
0
2
4
6
8
10
ΔK (MPa√m) 100 4 10
10
5
6
10
10
7
8
10
Cycles to Failure (N f)
PM 10
Fig. 8. Comparison of S–N curves for COMRAL-85 (R = 1).
and PM 20
p
unreinforced alloy p (3.50 ± 0.50 MPa m compared with 5.60 ± 1.97 MPa m). There was significant scatter in the results for the unreinforced alloy while the data for the liquid metallurgy processed composite were relatively consistent. The fatigue crack growth thresholds of the powder metallurgy processed composites are shown in Figs. 11–13 and a graph of DKth versus particle volume fraction is given in Fig. 14. The threshold values of these composites were also substantially lower than that of the unreinforced alloy.
AA 6061
Specimen No.
4
3
2
4
Specimen No.
TM
3
Threshold Value: 3.24 ± 0.41 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 11. Fatigue thresholds for PM 5 and PM 10 (R = 1).
However, there was no significant difference between the different volume fractions. The average values of DKth for the powder metallurgy processed and liquid metallurgy processed composites with the same volume fraction of reinforcement were the same, although the latter showed more scatter.
Threshold Value: 5.60 ± 1.97 MPa√m
3.3. Fractography 1 Solid Symbol: Unbroken Open Symbol: Broken
0 0
2
4
6
8
10
12
ΔK (MPa√m)
Fig. 9. Fatigue threshold for Al 6061-T6 (R = 1).
14
Fig. 15 shows the fatigue fracture surfaces of a number of the powder metallurgy processed composites at low magnification. For the 5% volume fraction composite, the fatigue fracture area was clearly visible on the fracture surface but there was no distinct fatigue fracture area in the
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
1263
9
PM 15
8
4
ΔKth (MPa√m)
Specimen No.
7 3 Threshold Value: 2.83 MPa√m
2
6 5 4 3 2
1 Solid Symbol: Unbroken Open Symbol: Broken
1 0
0
1
2
3
4
5
6
7
8
9
10
0
5
10
15
20
25
30
Particle Volume Fraction (%)
ΔK (MPa√m)
Fig. 14. Fatigue threshold versus particle volume fraction (R = 1). PM 20
Specimen No.
4
3
Threshold Value: 3.49 ± 0.29 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 12. Fatigue thresholds for PM 15 and PM 20 (R = 1).
Fig. 16 shows the microstructural features across the fatigue fracture surface of the powder metallurgy processed composite reinforced with 20% volume fraction of microspheres. The surface shows different microstructural features as DK increases. At near-threshold and low DK ranges, the surface was relatively featureless and flat. Only a few particles were revealed in this area. As DK increased, surface roughness was increased and more particles were revealed on the fracture surface. For higher DK, the surfaces were much rougher and similar to those of a tensile overload fracture. No fatigue striations were detected, as was also the case for all of the other composites investigated in this study.
PM 25
3.4. Crack propagation
Specimen No.
4
3
Threshold Value: 3.49± 0.29 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 17 shows a crack profile for the liquid metallurgy processed composite. Broken and debonded particles could be seen along the crack path. However, when the number of particles encountered on the crack path was compared with the number encountered on an arbitrary path, the number of particles associated with the crack profile was found to be considerably less than that encountered on the arbitrary path, Fig. 18. As shown in the micrograph of the crack tip area at high magnification, Fig. 17b, microvoids can be seen at the particle/matrix interface.
Specimen No.
PM 30
4
4. Discussion
3
4.1. Stress-life (S–N) behaviour Threshold Value: 2.83 MPa√m
2
1 Solid Symbol: Unbroken Open Symbol: Broken
0
1
2
3
4
5
6
7
8
9
10
ΔK (MPa√m)
Fig. 13. Fatigue thresholds for PM 25 and PM 30 (R = 1).
composites with microspheres.
the
higher
volume
fractions
of
It is generally accepted that the fatigue lives of PMMCs are longer than those of unreinforced alloys when stresscontrolled fatigue tests are performed [6–11]. The main contribution to the increased fatigue life in PMMCs is the higher elastic modulus. The fatigue process is strongly dependent on the strain amplitude, and the higher elastic modulus in PMMCs results in lower total strain, and especially total plastic strain, at any given stress level [12]. However, the improvement in fatigue life is not the same for all stress levels. As shown in Fig. 7, the increase in fatigue strength for the powder metallurgy processed composites
1264
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Fig. 15. Fatigue fracture surfaces at low magnification (Dr = 300 MPa, R = 0.2): (a) PM 5, (b) PM 20 and (c) PM 30.
Fig. 16. Fatigue fracture surface of PM 20 (Dr = 240 MPa, R = 0.2).
was not significant (or non-existent) at high stress levels but became increasingly more pronounced as the applied fluctuating stress decreased. This is similar to the behaviour observed by Vyletel et al. [7] in 15 vol.% TiC reinforced AA 2124. In fatigue tests under fully reversed loading (R = 1), fatigue is controlled by the plastic strain in the high-stress low-cycle fatigue regime (Nf < 104) while it is controlled by the elastic strain in the low-stress high-cycle regime (Nf > 104) [13]. Vyletel et al. [7] have suggested that, at high stress levels, the improvement in fatigue life expected in the composite as a result of the decreased plastic strain was offset by a decrease in fatigue life due to cyclic ductility, so that no net improvement occurs. As the stress level decreases, plastic behaviour becomes less important. The elastic behaviour now becomes increasingly dominant
and the fatigue life of the composite shows a progressive improvement over that of the unreinforced alloy since the elastic strain amplitude is less in the composite because of the higher elastic modulus. As discussed above, the elastic modulus is considered to have an important effect on fatigue life, at least in the highcycle low-stress regime. It is surprising therefore that, while the addition of particulate to the unreinforced alloy increased its fatigue life markedly, there was no further improvement with increasing volume fraction, even though a progressive increase in modulus occurred [14,15]. This finding is, however, consistent with the results obtained by Masuda and Tanaka [16] in their study of fatigue in A356 aluminium alloy reinforced with 10% and 20% SiC particles. They again found that the fatigue strength was
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
1265
Fig. 17. Fatigue crack profile for COMRAL-85TM: (a) high magnification of the rectangular area and (b) crack tip area.
higher in the composites than in the unreinforced alloy but observed no further improvement with increasing volume fraction. An absence of improvement in fatigue strength with increasing volume fraction has also been reported by Hall et al. [9] from their study of AA 2124 reinforced with 20% and 35% SiC particles. They found that the two composites also had similar tensile strengths and attributed the lack of improvement in fatigue life to the beneficial effect of increased modulus being offset by a degradation due to particle fracture. This would also explain the behaviour observed in the present study. Both static strength [14,15] and fatigue strength were not increased with increasing volume fraction of particulate. The former is considered to be due to particle fracture and this could also explain the lack of improvement in fatigue performance. A number of workers have attributed improvements in fatigue performance to increased static strength [16–19]. The findings of the present study show a relationship between fatigue performance and static strength [14,15] and are thus consistent with this view. However, it is unclear why the fatigue performances of PM 10 and PM 25 are inferior to the rest. The liquid metallurgy processed composite showed a shorter fatigue life than the powder metallurgy processed composite containing the same volume fraction of particulate. Moreover, the fatigue life of the liquid metallurgy processed composite was shorter even than that of the unreinforced alloy. This difference in fatigue life between the liquid metallurgy and powder metallurgy processed composites is not surprising in view of the different processing routes. Material manufactured by the powder metallurgy process is expected to have better fatigue strength
than that manufactured by the liquid metallurgy process because of the more refined grain size [20]. Indeed, Turnbull and Rios [21] have shown that in unreinforced AA 5754 the endurance stress corresponding to N = 107 load cycles was proportional to the inverse square root of the grain size. The grain size of the powder metallurgy processed composites used in the present study was definitely finer than that of the liquid metallurgy processed composite, and this difference in grain size is considered to be responsible, at least in part, for the difference in fatigue behaviour. Differences in grain size may also be responsible for some of the difference in fatigue life observed between the powder metallurgy processed composites and the unreinforced alloy. Another possible reason for the inferior fatigue performance of the liquid metallurgy processed composite is the more inhomogeneous distribution of the reinforcing particles (see Fig. 1). Yu et al. [22] found that the fatigue strength of AA 6061 reinforced with SiC particles was lower than that of the unreinforced alloy and attributed this to particle clustering. Similarly, the poor fatigue performance of the liquid metallurgy processed composite compared with that of the unreinforced alloy may be a result of the inhomogeneous particle distribution in the former. Ref. [23] contains a comprehensive summary of highcycle S–N data mostly with R = 0.1 (but also R = 1) for various aluminium alloys discontinuously reinforced with alumina or silicon carbide. The only study directly relevant to this work is by Lloyd [24] and contains stress amplitude versus cycles to failure curves at R = 1 for extruded unreinforced 6061-T6, and 6061-T6 reinforced with 15 vol.% Al2O3 particles. The S–N curves for
1266
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
6061-T6 and the composite agree moderately well with our data at low stress/high cycles to failure but the shapes of the curves differ, especially for high stress/low cycles to failure, where Lloyd’s curves diverge while ours cross.
increasing volume fraction is due to particle fracture. It is noted that some void formation was observed at the particle/matrix interface in front of growing fatigue cracks and some contribution from this effect may also have occurred.
4.2. Fatigue crack initiation
4.3. Fatigue crack growth threshold
Fatigue life consists of two components: crack initiation life and crack growth life. In smooth specimen tests, the fatigue crack initiation life can occupy a significant part of the total fatigue life. In un-notched specimens of high static strength aluminium alloys, crack initiation usually occupies between 75% and 95% of the total fatigue life [25]. The proportion of life consumed by crack initiation also increases with increasing static tensile strength. PMMCs have much higher tensile strengths than unreinforced alloys and their fatigue life would therefore be expected to be controlled by fatigue crack initiation. These conclusions are supported by the findings of several researchers [26]. In the case of unreinforced aluminium alloys, the main crack initiation mechanism is considered to be the formation of persistent slip bands (PSBs) [27]. In precipitation hardened aluminium alloys, fatigue life is controlled by the growth kinetics of PSBs in the low-cyclic load range. It is generally accepted that the material within the PSB lacks strengthening precipitates and is much softer than the surrounding matrix. However, PMMCs have more diverse crack initiation sites which are due to the particle additions. Srivatsan et al. [28] suggested that the particle/matrix interface is an important crack initiation site. During cyclic deformation the mismatch that exists between the hard and brittle reinforcing particles and the ductile matrix favours the concentration of stress at and near the particle/matrix interface, causing the matrix in the immediate vicinity to fail prematurely or the particle to separate from the matrix. On the other hand, Llorca and Poza [29] concluded that the dominant damage mechanism under cyclic deformation was reinforcement fracture, particles being broken by cracks perpendicular to the loading axis. Besides this kind of intrinsic crack initiation site, there may be extrinsic crack initiation sites such as particle clusters, casting defects (in ingot metallurgy processed composites), and pores. The propensity for particle fracture depends on the particle strength and size. Large and weak particles have a greater probability of fracturing than small and strong particles [30]. The MICRAL-20TM reinforcement particles used in this study have a broad size range [3] from 1 to about 50 lm with a Weibull modulus of 15 and the particles, being polycrystalline, appeared to be inherently weak. Substantial particle fracture appeared to occur below the yield strength of the composites [14,15] and it is most likely that particle fracture made a major contribution to fatigue crack initiation. This is consistent with the conclusion that the lack of improvement observed in fatigue life with
Previous studies of the effect of particulate addition on fatigue threshold have given mixed results with some authors reporting an increase in DKth [31] and others reporting a decrease [32], as was found here. This contradictory behaviour appears to be due to differences in the material and in the testing technique. There are some important differences between the method used here and that used by others. In the present work, the material was obtained as small diameter extruded bar and this precluded the use of compact tension specimens, which have been used in most other work where the material has been obtained as plate. In this case, round specimens were used instead. In addition, the tests were performed at R = 1 whereas most other workers have used a low value of R (generally R < 0.3). It is significant, however,pthat Crawford [33] obtained a threshold value of 3.6 MPa m for the 20 vol.% liquid metallurgy processed composite COMRAL-85TMpwhich is in excellent agreement with the value of 3.5 MPa m obtained in the present study. Moreover, this result was obtained for R = 0.1 using a compact tension test specimen. This may indicate that the threshold behaviour is very dependent on the specific composite. The absence of any systematic change in DKth with volume fraction is surprising in view of the generally held belief that volume fraction is one of the parameters which influence fatigue threshold [34]. The results are, however, consistent with the findings of Sugimura and Suresh [32] who reported an absence of any significant change in DKth with volume fraction in Al–3.5Cu composites reinforced with SiC particles. It is noted that, as in the present study, DKth of the composite was found to be lower than that of the unreinforced alloy. In the present work, no difference was found in DKth between the cast and powder metallurgy processed composites, yet again differences are commonly reported [35]. These differences have been attributed to microstructural differences resulting in enhanced roughness-induced crack closure. However, when fatigue testing is performed under tension–compression conditions, that is R = 1, as in the present study, the crack closure effect is negligible. Kumai et al. [35] have confirmed that when the crack closure effect is removed, the DKth of cast composites is the same as that of powder metallurgy processed composites as found in the present study. Ref. [23] contains a comprehensive summary of fatigue crack growth data mostly with R = 0.1 (but also R = 1) for various aluminium alloys discontinuously reinforced with alumina or silicon carbide. The only study directly relevant to this work is by Lloyd [24] and contains fatigue
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
crack growth rate curves at R = 1 for 6061-T6, and 6061T6 reinforced with 15 vol.% Al2O3 particles. The threshold stress intensity range DKth for 6061-T6 is about p p 4.3 MPa m in Lloyd’s paper versus 5.60 ± 1.97 MPa m in our work, while p that for the 15% Al2O3 composite p is about 6.1 MPa m in Lloyd’s work versus 2.83 MPa m in ours. This higher threshold in Lloyd’s work is (wrongly) attributed to crack closure effects. 4.4. Fatigue crack propagation
Cumulative Number of Particles
As shown in Fig. 16, only a few particles were revealed on the fracture surface at the near-threshold and low DK ranges. This indicates that in the early stage of crack growth, the fatigue crack tends to avoid the particles and propagate through the matrix. This was also apparent from the measurements made along the crack profile which indicated that the crack propagated through a smaller number of particles than would be encountered along an arbitrary line (refer to Fig. 18). However as DK increased, more particles became evident on the fracture surface, indicating that the crack was starting to seek out fractured and debonded particles in front of the growing crack tip (see Figs. 16 and 17). The lack of fatigue striations on the surface at the mid-range of DK suggested that the fatigue crack propagated in a predominantly brittle manner. These findings are consistent with the results obtained by Kumai et al. [36] in their study of short and long fatigue crack propagation in AA 6061 alloy reinforced with 15 vol.% of SiC particles. They concluded that at low stress levels, the fatigue crack tended to avoid SiC particles and ran through the matrix. As the stress level increased, the particles began to crack and the fatigue crack then linked up with the cracked particles. Fatigue crack propagation is strongly dependent on the interaction between particles and the advancing crack tip. As shown by Padkin et al. [37], when the modulus of second phase particles in front of the advancing fatigue crack is higher than that of the matrix alloy, as in the present study, the crack tends to avoid the particles. However as 28 26 24 22 20 18 16 14 12 10 8 6 4 2 0
Random Path Crack Path
0
0.2
0.4
0.6
0.8
1
1.2
1.4
1.6
1.8
Distance from Notch (mm)
Fig. 18. Numbers of particles encountered on a random (circle symbol) and crack (square symbol) path.
1267
the crack grows, DK increases, so that the plastic zone size in front of the crack tip becomes larger and progressively more particles become included in the plastic zone. This increases the propensity for particle fracture and debonding. Some evidence of microvoid formation at the particle/ matrix interface was seen in the present study. This is consistent with in situ SEM observations of fatigue crack propagation in the liquid metallurgy processed composite COMRAL-85TM made by Hadianfard et al. [38] which showed that fatigue crack growth was strongly associated with the particle/matrix interface. 5. Conclusions The fatigue strength of the powder metallurgy processed composites was generally higher than that of the unreinforced alloy, as found in other studies on PMMCs. The improved performance of the composites is attributed to their higher elastic moduli. The beneficial effect of particle addition on fatigue strength was more evident at lower stress levels. There was no significant change in fatigue strength with increasing volume fraction of particulate. The absence of any effect of the particle volume fraction on fatigue strength is considered to be due to the beneficial effect of the increase in modulus with volume fraction being offset by a degradation due to particle fracture. The fatigue strength of the liquid metallurgy processed composite was lower than those of the powder metallurgy processed composites and also less than that of the unreinforced alloy. The poorer performance of the liquid metallurgy processed composite compared with the powder metallurgy processed composites is attributed to its coarser grain size. The DKth values for the composites were lower than for the unreinforced alloy, in agreement with the results of some prior studies but contrary to the results of others. There was no systematic change in the DKth value with particle volume fraction as found previously in SiC particle reinforced Al–3.5Cu. The DKth values for the powder metallurgy and liquid metallurgy processed composites with the same volume fraction (20%) of reinforcement were similar. This is attributed to the microstructural differences between the two types of composite becoming insignificant for the R value of 1 used in the present study. Growing fatigue cracks were found to avoid the particles, particularly at lower DK values, whereas growing cracks produced during static fracture were attracted to the particles. This difference in behaviour appears to be due to a difference in the process zone size. While the process zone is small, as in the early stages of fatigue, the growing crack is deflected by the particles due to their high modulus. As the process zone becomes larger, it incorporates progressively more particles and the propensity for particle fracture and decohesion increases.
1268
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269
Acknowledgements This work was supported by Comalco Research Centre (15 Edgars Road, Thomastown, VIC 3074, Australia) with additional financial assistance through an Australian Research Council small grant. The assistance and advice provided by Drs. M.J. Couper and K. Xia of CRC are gratefully acknowledged. Mr. R.P. Harrison of the Australian Nuclear Science and Technology Organisation (ANSTO) is thanked for the loan of a Schenck horizontal midget pulser fatigue testing machine. References [1] Couper MJ, Xia K. Development of microsphere reinforced metal matrix composites. In: Hansen N, Juul Jensen D, Leffers T, Lilholt H, Lorentzen T, Pedersen AS, Pedersen OB, editors. Proceedings of the twelfth Risø international symposium on metallurgy and materials science: metal matrix composites – processing, microstructure and properties. Risø National Laboratory, Roskilde, Denmark, September; 1991. p. 291–8. [2] Xia K. Interfaces and interfacial reactions in Comral composites. In: Bandyopadhyay S, Crosky AG, editors. Proceedings of the second Australian forum on metal matrix composites. University of New South Wales, Sydney, Australia, December; 1991. p. 1–2. [3] Comalco data sheets. Comalco Research Centre, Thomastown, Victoria, Australia; 1990. [4] ASTM E466-82, Standard practice for conducting constant amplitude axial fatigue tests of metallic materials. Philadelphia: American Society for Testing and Materials; 1982. [5] Murakami Y. Stress intensity factors handbook, 1st ed. vol. 1. Oxford: Pergamon; 1987. p. 40. [6] Bonnen JJ, Allison JE, Jones JW. Fatigue behavior of a 2xxx series aluminum alloy reinforced with 15 vol pct SiCp. Metall Trans A 1991;22A(5):1007–19. [7] Vyletel GM, Allison JE, Van Aken DC. The influence of matrix microstructure and TiC reinforcement on the cyclic response and fatigue behavior of 2219 Al. Metall Trans A 1993;24A(11):2545–57. [8] Harris SJ, Dinsdale K, Gao Y, Noble B. Influence of heat treatment on the monotonic and fatigue properties of aluminium alloy composites. In: Andersen SI, Lilholt H, Pedersen OB, editors. Proceedings of the ninth Risø international symposium on metallurgy and materials science: mechanical and physical behaviour of metallic and ceramic composites. Risø National Laboratory, Roskilde, Denmark, September; 1988. p. 373–82. [9] Hall JN, Jones JW, Sachdev AK. Particle size, volume fraction and matrix strength effects on fatigue behavior and particle fracture in 2124 aluminum–SiCp composites. Mater Sci Eng A 1994;A183(1–2): 69–80. [10] Liu J, Wang DZ, Yao CK, Yu WC, Wang ZG. Effects of ageing treatment on fatigue strength in a SiCw/6061 Al composite. J Mater Sci Lett 1992;11(14):970–2. [11] Shang JK, Ritchie RO. Fatigue of discontinuously reinforced metal matrix composites. In: Everett RK, Arsenault RJ, editors. Metal matrix composites. San Diego: Academic Press; 1991. p. 255–85. [12] Allison JE, Jones JW. Fatigue behaviour of discontinuously reinforced metal matrix composites. In: Suresh S, Mortensen A, Needleman A, editors. Fundamentals of metal–matrix composites. Boston: Butterworth; 1993. p. 275 [Chapter 15]. [13] Barsom JM, Rolfe ST. Fracture and fatigue control in structures: applications of fracture mechanics. 2nd ed. Englewood Cliffs, New Jersey: Prentice-Hall; 1987. p. 239. [14] Park BG, Crosky AG, Hellier AK. Material characterisation and mechanical properties of Al2O3–Al metal matrix composites. J Mater Sci 2001;36(10):2417–26.
[15] Park BG. Fracture and fatigue of Al2O3 particulate reinforced metal matrix composites. PhD thesis. University of New South Wales, Sydney, Australia; 1997. [16] Masuda C, Tanaka Y. Tensile and fatigue properties of SiC whiskers and SiC particulate reinforced aluminium composites. In: Proceedings of the seventh international conference on fracture (ICF7): advances in fracture research, vol. 4. Houston, Texas, USA, March; 1989. p. 3089–96. [17] Hurd NJ. Fatigue performance of alumina reinforced metal matrix composites. Mater Sci Technol 1988;4(6):513–7. [18] Mura T, Qin S, Fan H. A theory of fatigue crack initiation. In: Teoh SH, Lee KH, editors. Proceedings of the joint FEFG/ICF international conference on fracture of engineering materials and structures. Singapore, August, 1991. London and New York: Elsevier Applied Science; 1991. p. 57–68. [19] Harris SJ, Dinsdale K, Gao Y, Noble B. Influence of heat treatment on the monotonic and fatigue properties of aluminium alloy composites. In: Andersen SI, Lilholt H, Pedersen OB, editors. Proceedings of the ninth Risø international symposium on metallurgy and materials science: mechanical and physical behaviour of metallic and ceramic composites. Risø National Laboratory, Roskilde, Denmark, September; 1988. p. 378. [20] Allison JE, Jones JW. Fatigue behaviour of discontinuously reinforced metal–matrix composites. In: Suresh S, Mortensen A, Needleman A, editors. Fundamentals of metal–matrix composites. Boston: Butterworth; 1993. p. 269 [Chapter 15]. [21] Turnbull A, de los Rios ER. The effect of grain size on fatigue crack growth in an aluminium magnesium alloy. Fatigue Fract Eng Mater Struct 1995;18(11):1355–66. [22] Yu S-Y, Ishii H, Tohgo K. Corrosion fatigue of SiC whisker or SiC particulate reinforced 6061 aluminum alloy. Fatigue Fract Eng Mater Struct 1994;17(5):571–8. [23] Lewandowski JJ. Fracture and fatigue of particulate MMCs. In: Kelly A, Zweben C, editors-in-chief. Comprehensive composite materials, vol. 3. Oxford: Pergamon; 2000. p. 151–87 [Section 3.07]. [24] Lloyd DJ. Particle reinforced aluminium and magnesium matrix composites. Int Mater Rev 1994;39(1):1–23. [25] Hahn CT, Simon R. A review of fatigue crack growth in high strength aluminum alloys and the relevant metallurgical factors. Eng Fract Mech 1973;5(3):523–40. [26] Wang DZ, Liu J, Yao CK, He ZH, Geng L. The initiation of fatigue microcracks near SiC whisker ends in a silicon carbide whisker reinforced 6061 aluminium composite. J Mater Sci Lett 1994;13(13):957–9. [27] Baxter WJ, McKinney TR. Growth of slip bands during fatigue of 6061-T6 aluminum. Metall Trans A 1988;19A(1):83–91. [28] Srivatsan TS, Auradkar R, Prakash A, Lavernia EJ. Influence of microstructure on cyclic stress response and cyclic fracture behavior of aluminium alloy metal–matrix composite. In: Teoh SH, Lee KH, editors. Proceedings of the joint FEFG/ICF international conference on fracture of engineering materials and structures. Singapore, August, 1991. London and New York: Elsevier Applied Science; 1991. p. 194–202. [29] Llorca J, Poza P. Influence of reinforcement fracture on the cyclic stress–strain curve of metal–matrix composites. Acta Metall Mater 1995;43(11):3959–69. [30] Curry DA, Knott JF. Effects of microstructure on cleavage fracture stress in steel. Metal Sci 1978;12(11):511–4. [31] Llorca J, Ruiz J, Healy JC, Elices M, Beevers CJ. Fatigue crack propagation in salt water, air and high vacuum in a spray-formed particulate-reinforced metal matrix composite. Mater Sci Eng A 1994;A185(1–2):1–15. [32] Sugimura Y, Suresh S. Effects of SiC content on fatigue crack growth in aluminum alloys reinforced with SiC particles. Metall Trans A 1992;23A(8):2231–42. [33] Crawford BR. Fatigue of an aluminium-6061 matrix composite. PhD thesis. University of Queensland, St. Lucia, Australia; 1996.
B.G. Park et al. / Composites: Part B 39 (2008) 1257–1269 [34] Davidson DL. The growth of fatigue cracks through particulate SiC reinforced aluminum alloys. Eng Fract Mech 1989;33(6):965–77. [35] Kumai S, Hu J, Higo Y, Nunomura S. Effects of dendrite cell size and particle distribution on the near-threshold fatigue crack growth behaviour of cast Al–SiCp composites. Acta Mater 1996;44(6): 2249–57. [36] Kumai S, King JE, Knott JF. Short and long fatigue crack growth in a SiC reinforced aluminium alloy. Fatigue Fract Eng Mater Struct 1990;13(5):511–24.
1269
[37] Padkin AJ, Brereton MF, Plumbridge WJ. Fatigue crack growth in two-phase alloys. Mater Sci Technol 1987;3(3):217–23. [38] Hadianfard MJ, Healy JC, Mai Y-W. The influence of heat treatment on mechanical and fracture properties of an alumina microsphere reinforced aluminium matrix composite. In: Chandra T, Dhingra AK, editors. Proceedings of advanced composites ’93. University of Wollongong, Australia, February; 1993. p. 1199–1206.