Creep properties of an Al-2024 composite reinforced with SiC particulates

Materials Science and Engineering A328 (2002) 39 – 47 www.elsevier.com/locate/msea

Creep properties of an Al-2024 composite reinforced with SiC particulates Stefano Spigarelli a,*, Marcello Cabibbo a, Enrico Evangelista a, Terence G. Langdon b b

a INFM/Department of Mechanics, Uni6ersity of Ancona, I-60131 Ancona, Italy Departments of Aerospace and Mechanical Engineering and Materials Science, Uni6ersity of Southern California, Los Angeles, CA 90089 -1453, USA

Received 19 February 2001; received in revised form 28 May 2001

Abstract Creep tests were conducted on an Al-2024 alloy reinforced with 15% SiC particulates. In this composite, as in earlier experiments on the unreinforced Al-2024 alloy, there is a continuous precipitation of fine particles throughout the tests. The creep behaviour of the composite reveals the presence of a threshold stress, and this stress depends upon both temperature and applied stress because more copious precipitation occurs in the more time-consuming tests conducted at the lower stress levels. The results for the composite are analysed using a procedure similar to that developed earlier for the unreinforced alloy. It is shown that the rate-controlling deformation mechanism is dislocation climb within the matrix alloy with a stress exponent of about 4.4 and an activation energy equal to that anticipated for lattice self-diffusion in pure aluminium. © 2002 Elsevier Science B.V. All rights reserved. Keywords: Aluminium alloys; Creep; Composites; Threshold stress

1. Introduction The creep behaviour of pure metals and solid solution alloys is now understood reasonably well. The creep curves of strain against time usually exhibit welldefined steady-state conditions and logarithmic plots of the measured steady-state creep rate, m; s, against the applied stress, |, lead to a series of straight lines for each testing temperature with the slopes of the lines delineating the stress exponent, n, which lies within the range of  3–6. The creep behaviour is often different in metal matrix composites (MMCs) and in some alloys containing precipitates where the secondary stage of creep may be sufficiently short that it is more reasonably described as a minimum creep rate, m; m [1]. When creep data for MMCs are collected over a sufficiently wide range of strain rates, logarithmic plots of either m; s or m; m against | often reveal a markedly downward curvature so that the value of n increases sharply at the * Corresponding author. Tel.: + 39-071-220-4746; fax: + 39-071220-4799. E-mail address: [email protected] (S. Spigarelli).

lowest levels of the applied stress. Under these conditions, it is generally assumed that deformation occurs under the action of an effective stress, |e, which is defined as |− |0 where |0 is a threshold stress delineating the lower limiting stress for measurable flow [2]. The minimum creep rate is then expressed by a relationship of the form



DGb | − |0 m; m = A kT G



n

,

(1)

where D is the appropriate diffusion coefficient, G is the shear modulus, b is the magnitude of Burgers vector, k is Boltzmann’s constant, T is the absolute temperature and A is a dimensionless constant. The presence of a threshold stress, and the consequent use of Eq. (1), is not restricted solely to the creep of MMCs. In a detailed investigation of the creep behaviour of an unreinforced Al-6061 alloy produced by powder metallurgy (PM), it was shown that the value of n increased sharply at the lowest stress levels and it was necessary to incorporate a threshold stress into the analysis due, it was suggested, to the interaction of dislocations with fine incoherent oxide particles

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introduced during the atomization stage in the PM processing [3]. Similar creep behaviour, including the use of Eq. (1), was reported also for unreinforced PM Al-2124 [4] and PM Al-6092 [5] alloys. A recent report described the high temperature creep properties of an unreinforced PM Al-2024 alloy [6]. This material also exhibited behaviour, where it was necessary to invoke the presence of a threshold stress but the experimental results were unusual because, although large particles of Al2Cu and Al2CuMg were present prior to creep testing, there was also a continuous precipitation of very fine ( 60 nm) particles throughout the creep tests. This behaviour was successfully analysed by introducing a time-dependent relationship for the precipitation process [7] so that the Table 1 Chemical compositions (in wt.%) for the matrix alloy of the composite and the unreinforced Al-2024 alloy tested earlier [6,7]

Composite matrix Unreinforced alloy

Si

Fe

Cu

Mg

Mn

Al

0.01 0.24

0.03 0.26

3.99 4.42

1.60 1.47

0.61 0.56

Bal. Bal.

creep data yielded a stress exponent of  5 and an activation energy close to the value anticipated for lattice self-diffusion in pure aluminium [6]. To date, there appear to be no detailed reports of the creep behaviour of an MMC where there is a continuous precipitation throughout the creep life of the material. Accordingly, the present experiments were initiated to evaluate the creep properties of an MMC having PM Al-2024 as the matrix alloy and with a reinforcement of SiC particulates. As will be demonstrated, the creep characteristics of this material have similarities to those reported earlier for the unreinforced PM Al-2024 alloy [6] and it is possible to interpret the data using a similar analytical procedure.

2. Experimental materials and procedures The experiments were conducted on an Al-2024 composite reinforced with 15 vol. of irregularly shaped SiC particulates. The composite was produced by PM with the powders compacted and subsequently sintered at a temperature of 773 K and then extruded at 723 K followed by air cooling to room temperature. The powders were atomized under inert gas and their chemical composition is given in Table 1. Cylindrical specimens were machined from the extruded bars, solution-treated at 673 K for 5 h and then slowly cooled at  60 K h − 1. The unreinforced PM Al-2024 alloy was tested in the earlier investigation after similar processing and solution treatment [6]: for convenience, the chemical composition of the unreinforced alloy is also included in Table 1. Inspection showed the composite contained an array of essentially equiaxed grains prior to creep testing and these grains were very fine with an average size of the order of 3 mm: this grain size is similar to that reported for the unreinforced alloy [6]. All tensile creep tests were conducted in air under conditions of constant load and at absolute temperatures in the range from 548 to 603 K. Following creep testing, the microstructures of several samples were examined using transmission electron microscopy.

3. Experimental results

3.1. Nature of the creep cur6es

Fig. 1. (a) Strain rate vs strain for three samples tested at 573 K. (b) Comparison between a constant load test and a stress-change test.

The creep curves obtained in these experiments revealed a short primary stage, a minimum in the creep rate and then a very extended tertiary stage. These trends are readily apparent in Fig. 1(a) where the instantaneous strain rate, m; , is plotted against the measured strain for representative tests conducted at 573 K under three different levels of the applied stress. The appearance of these curves is essentially identical to

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Fig. 1(b) shows the effect, at a temperature of 603 K, of a change in stress from 30 to 38 MPa compared with a test conducted under a constant load of 38 MPa. It is apparent from these two curves that the increase in the time of creep exposure leads, after the stress change, to a substantially lower minimum creep rate. It is therefore reasonable to conclude that creep exposure is accompanied by some form of strengthening, probably related to the precipitation of fine particles that obstruct dislocation mobility. Fig. 2(a) shows a logarithmic plot of the minimum creep rate, m; m, against the applied stress, |, for samples tested at three different temperatures: for convenience, similar data for the unreinforced Al-2024 alloy [6] are given in Fig. 2(b) for the same three testing temperatures. Inspection of Fig. 2 shows that, for both the composite and the unreinforced alloy, the data extend over  5 orders of magnitude of strain rate and the curves give high values of n, up to n: 15, at the lowest strain rates. This suggests the presence of a threshold stress, |0, in both materials. A simple procedure for estimating the values of |0 at each testing temperature is to extrapolate the data to very low strain rates so that the experimental lines become vertical [10]. If this simple procedure is used with the data in Fig. 2(a) and (b), it is apparent that the values of |0 are slightly higher in the MMC than in the unreinforced alloy.

3.2. Characteristics of the microstructures after creep testing

Fig. 2. Minimum creep rate vs applied stress for (a) the composite material and (b) the unreinforced alloy [6].

Figs. 3 and 4 show examples of the microstructures after creep testing at the highest temperature of 603 K and at stresses of 45 and 32 MPa: these two conditions correspond to total testing times of 1.45× 103 and 9.7× 105 s, respectively. In general, careful inspection

Fig. 3. Microstructure after testing at 603 K with an applied stress of 45 MPa: total testing time =1.45× 103 s. Very fine precipitates decorating dislocations can be observed.

that reported earlier for the unreinforced Al-2024 alloy [6] and it is also similar to plots of m; vs strain reported for MMCs fabricated using both PM [8] and liquid metallurgy [9] procedures.

Fig. 4. Microstructure after testing at 603 K with an applied stress of 32 MPa: total testing time =9.76 × 105 s. Comparison with Fig. 3 clearly suggests that precipitation proceeded during high temperature exposure, leading to an increase in volume fraction and size of the precipitates.

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of a number of samples revealed similarities with the observations recorded earlier for the unreinforced Al2024 alloy [7]. Specifically, there was a distribution of relatively large (450 nm) Al2Cu and Al2CuMg particles on the grain boundaries and within the grains prior to testing and, in addition, there was a continuous precipitation of fine (10 – 200 nm) particles within the grains during creep. It was found that, as in the unreinforced alloy [7], long-term creep testing at all temperatures from 548 to 603 K led to a very significant increase in the density of these fine precipitates. Fig. 3, in particular, shows extremely fine precipitates, 10 nm or less in diameter. A longer time of exposure, as in the case of Fig. 4, results in an increase in volume fraction of these precipitates, and in a parallel increase in their size.

4. Discussion

4.1. The stress dependence of creep The data in Fig. 2(a) are consistent with the presence of a threshold stress and therefore with the use of Eq. (1). Furthermore, the presence of reasonably similar threshold stresses in both the composite and the unreinforced alloy suggest the origin of |0 may be related, at least in part, to the continuous precipitation of the fine precipitates. It was shown earlier that the precipitation kinetics in the unreinforced Al-2024 alloy may be expressed by a relationship of the form [7] Nv = Nv0 +DNv[1−exp( −q/t%)],

(2)

where Nv is the instantaneous number of fine particles per unit volume, Nv0 is the initial number of fine particles per unit volume, DNv is the total additional density of fine particles produced by precipitation during creep, q is a temperature-compensated time (= t exp( − Q/RT), where t is the total time of testing, Q is the appropriate activation energy associated with the precipitation process and R is the gas constant) and t% is a time parameter. Observations suggested that the continuous precipitation occurring during the creep of the composite in this investigation was qualitatively similar to that reported earlier for the unreinforced Al-2024 alloy [7]. Therefore, since the matrix alloy of the composite is Al-2024, and both the reinforced and the unreinforced materials were creep tested over the same ranges of stress and temperature, it is appropriate to make use of the analytical procedure developed earlier for the Al-2024 alloy. The rate of precipitation in the matrix alloy is controlled by lattice self-diffusion [7] with an activation energy is equal to the value for self-diffusion in pure Al (Q :143.4 kJ mol − 1 [11]). It is reasonable to assume that, after an infinite testing time, all of the Mg is

precipitated in the form of Al2CuMg particles with the remaining Cu precipitated as Al2Cu. From an earlier calculation based on the chemical composition of the matrix alloy [7], it was shown that DNv : 1.6×1020 m − 3 and, from a best fit to experimental measurements for the unreinforced Al-2024 alloy using Eq. (2), Nv0 : 7.8× 1018 m − 3 and t% :1.1× 10 − 6 s. The precise nature of the threshold stress in the unreinforced alloy and in the composite is dependent upon the method by which the dislocations overcome local obstacles impeding their forward movement. These local obstacles are in the form of both large ( 450 nm) precipitates present prior to creep testing and fine precipitates continuously precipitated throughout each test. Three possible mechanisms may be identified for overcoming these obstacles and these are associated with the Orowan stress, |Or, required to bow the dislocations between the particles [12,13], the back stress, |b, needed to create the additional dislocation segment as dislocations surmount the obstacle by local climb [14,15] and the detachment stress, |d, associated with detaching a dislocation from an attractive particle [16–18]. Analysis of data for the unreinforced Al-2024 alloy showed that, by taking creep results obtained at a temperature of 523 K where there was no significant precipitation of fine particles during creep, the measured threshold stress in the alloy was significantly larger than the values estimated for either |b or |d but it was in reasonable agreement with the value deduced for the Orowan stress, |Or [6]. The situation is complicated at creep testing temperatures above 523 K because the continuous precipitation of fine particles requires an estimate of the total extent of precipitation up to the occurrence of the minimum creep rate, m; m. It was shown for the unreinforced alloy that the measured threshold stresses at these higher temperatures represented a decreasing fraction of the Orowan stress as the temperature increased. The value of the Orowan stress, |Or, may be expressed as [12,13] |Or = 0.84M

Gb , (u−dp)

(3)

where M is the appropriate Taylor factor ( 3), G= {(3.022× 104)− (16.0×T)} MPa [11] and b=2.86× 10 − 10 m for pure Al, dp is the diameter of the fine particles (60 nm in the matrix alloy) and u is the interparticle spacing given by u=

  y 6fv

1/2

r,

(4)

where fv is the volume fraction of particles and r is the average particle radius. In practice, the term Nv in Eq. (2) is related to the interparticle spacing, u, through the expression [6]

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inforced alloy and the MMC. An alternative procedure was therefore adopted by expressing the threshold stress in the composite by a relationship of the form |0 = |0i + a exp(− s|)=h%|Or,

Fig. 5. Estimated variation of the threshold stress with the applied stress for the unreinforced alloy (solid lines) and the composite material (dashed lines).

Fig. 6. Minimum creep rate vs effective stress for the composite and the unreinforced alloy [6].

u:

1 2 Nvdp

.

(5)

The preceding relationships may be used to estimate the values of the threshold stress, |0, in the unreinforced alloy for conditions pertaining to the minimum creep rate for each separate testing condition. Furthermore, since the time of testing up to the strain associated with each minimum creep rate is dependent upon the magnitude of the stress used in each experiment, it follows that the values of the threshold stress, |0, are dependent also upon the level of the applied stress. This is illustrated by the solid lines for the unreinforced alloy in Fig. 5. In principle, the same approach may be used also to estimate the magnitudes of the threshold stresses for the composite but this necessitates assuming, a priori, that the precipitation kinetics are identical in both the unre-

(6)

where |0i is the threshold stress corresponding to the initial distribution of particles present in the material before creep testing, a and s are temperature-dependent terms which serve to incorporate the effect of the additional precipitation occurring up to the time of the minimum creep rate, and the final expression recognizes the proportionality between the threshold stress and the Orowan stress through a factor h% which also depends on temperature. Thus, the threshold stresses in the composite may be estimated by substituting the second expression of Eq. (6) into Eq. (1), making the reasonable assumption that the value of A in Eq. (1) is identical for the MMC and the unreinforced alloy, calculating the value of DGb/kT for each testing temperature, and then using a best fit procedure to estimate the values of a and s at each testing temperature. This calculation was performed by taking D equal to the value for lattice self-diffusion in aluminium (1.86× 10 − 4 exp(− 143.4/RT) m2 s − 1 [11]) and with a stress exponent of n= 4.4. As will be demonstrated, these values of D and n are consistent with the experimental data and a value of n=4.4 was selected because it is the stress exponent reported both for the creep of pure Al [19] and for some Al-based solid solution alloys [20] and MMCs with Al alloy matrices [21] when dislocation climb is the rate-controlling mechanism. From this analysis, the values of the threshold stresses were estimated for the composite and they are denoted by the dashed lines in Fig. 5 for the three different testing temperatures. Inspection of Fig. 5 leads to two conclusions. First, the values of |0 tend to increase at the lower levels of the applied stress, especially at the highest testing temperature of 603 K. This increase arises at the lower stresses because the total testing times to the advent of the minimum creep rate are then sufficiently long that many fine particles are introduced into the matrix through continuous precipitation. Second, although the trends for the MMC and the unreinforced alloy are similar, there are higher values of |0 in the composite for any selected testing temperature although the difference is generally less than a factor of 2. Finally, using the values estimated for |0 for each testing condition, Fig. 6 shows a plot of the minimum creep rate, m; m, against the effective stress, |− |0, for both the unreinforced Al-2024 alloy [6] and the composite reinforced with 15% of SiC particulates. It is apparent the datum points fall on separate lines for each testing temperature and each line has a slope close to 4.4 which is consistent with dislocation climb in the matrix alloy as the rate-controlling mechanism. It is

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apparent, therefore, that the creep behaviour of the composite is more similar to pure Al with n : 4.4 under climb conditions than some dilute Al– Mg alloys where n= 3.0 and dislocation glide is rate-controlling due to the dragging of solute atom atmospheres formed around the moving dislocations [22,23]. In practice, the evidence for control by dislocation climb is reasonable when it is noted that, although the Al-2024 matrix alloy contains 1.5 wt.% Mg implying a possibility of control by dislocation glide, there is a very severe depletion of the Mg in solid solution because of the formation of the Al2CuMg particles.

4.2. The temperature dependence of creep The activation energy for creep in the composite may be estimated by constructing a semi-logarithmic plot of m; mG n − 1T against 1/T for any selected value of |−|0. The result is shown in Fig. 7 for n=4.4 and an effective stress of 10 MPa giving, as expected, an activation energy of Q: 145 kJ mol − 1 which is in excellent agreement with the value anticipated for self-diffusion in aluminium ( 143.4 kJ mol − 1 [11]). Taking D for self-diffusion in pure Al, all of the datum points in Fig. 6 may now be replotted in a normalized form as m; mkT/DGb vs (| − |0)/G, as shown in Fig. 8 where all points scatter around a single line with a slope of n=4.4.

4.3. Shape of the creep cur6e and 6ariation in the threshold stress

Fig. 7. Semi-logarithmic plot of m; mG n − 1T vs 1/T at an effective stress of 10 MPa, giving an activation energy for creep of Q= 145 kJ mol − 1.

As clearly demonstrated in Fig. 1, the strain rate vs strain curves change in shape as the stress decreases. This behaviour deserves further examination because it is a consequence of two concurrent processes: the increase in the true applied stress and the coarsening of the precipitates. The former effect may be readily analysed by substituting into Eq. (1) the expression for the true applied stress, |%= |(1+ m), where, at a time corresponding to the minimum creep rate, |%:|. If |0 does not change with time and/or strain, this variation in the true applied stress leads to a substantial increase in the strain rate which is especially evident in the low stress regime where | is relatively close to the threshold stress. However, this explanation cannot account for the dramatic drop, followed by the sharp increase in strain rate, which is a characteristic feature of the low stress curves in Fig. 1(a). A similar behaviour has been reported also in other particle-strengthened alloys [24] and was attributed to massive precipitation followed by coarsening of the precipitates. A few simple calculations permit a semi-quantitative estimate of the increase in particle size as coarsening proceeds. For example, at 573 K with a stress of 38 MPa the value of |0 is calculated as 29 MPa corresponding to the minimum in creep rate. If it is assumed that no other softening mechanism or damage mechanism is operating, it follows from Eq. (1) that the value of the threshold stress generated by a dispersion of particles at a given strain (|0i) can be estimated from the relationship |0i = |−

Fig. 8. Temperature-compensated minimum creep rate vs normalized effective stress for the composite and the unreinforced alloy [6].

 m; i

m; m

1/n

(| −|0),

(7)

where m; i is the strain rate at m= mi and |0 again denotes the threshold stress corresponding to the minimum strain rate. This procedure gives, at m=0.02 (t=1.5 × 105 s), |02 : 24 MPa. Additionally, if |0i = h%|Or and

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the precipitation is complete so that fv is a constant in Eq. (4), a combination of Eqs. (3) and (4) gives |0 dpi = , |0i dp

(8)

where dp is again the particle size at the time corresponding to the minimum strain rate and dpi is the particle size at m =mi. Thus, in the case of the test carried out at 573 K with a stress of 38 MPa, an increase in strain rate of one order of magnitude can arise simply from a 20% increase in the average precipitate size due to coarsening. However, the prerequisites for this mechanism are a sufficiently high testing temperature and a long time of creep exposure so that the effect is observed only in the low stress regime at high temperatures. Additional evidence of the precipitation process can be obtained by considering the lowest stress data at 573 and 603 K; a careful inspection of Figs. 2(a) and 6 shows these data were not included in the latter plot. This omission is related to the very low, or even negative, values of the effective stress when the curves in Fig. 5 are used to estimate the threshold stresses under these conditions. It is obvious that Eq. (6) has an empirical nature since its scope is simply to interpolate the threshold stress values in a limited range of experimental conditions. A more physically sound relationship, Eq. (2), in combination with Eqs. (3)– (5), predicts that the threshold stress increases with time up to a maximum and thereafter it remains constant because the model fails to incorporate any effect of particle coarsening. In practice, the curve in Fig. 1(b) provides a direct indication on the magnitude of the threshold stress. Considering the minimum in the strain rate after the stress change so that m; m :3 ×10 − 8 s − 1, it follows from Fig. 6, for |=38 MPa, that the value of the threshold stress is close to 32 MPa. Thus, this value is larger than the initial applied stress (30 MPa) but lower than the value obtained by extrapolating the relevant curve in Fig. 5. The fact that creep occurs under these two separate conditions (603 K and 30 MPa or 573 K and 32 MPa) suggests that the particle by-passing mechanism changes when the applied stress is very close to the threshold stress. A similar behaviour was reported in another composite [25] where, in the very low stress regime at high temperatures, there was a marked inflection in the strain rate vs stress curves. After the stress change and the new minimum in strain rate, the curve in Fig. 1(b) exhibits a marked increase in strain rate that contrasts with the more moderate increase for the case of the constant load test. This is further evidence that the increase in strain rate normally observed in the very low stress tests, which is typical of the advent of the tertiary stage of creep, is not related exclusively to the magnitude of the applied stress but rather it is associated with the very long time of creep exposure at a high temperature.

45

Thus, it can be reasonably inferred that the threshold stress does not increase indefinitely as the empirical equation (6) predicts but rather the threshold stress reaches a maximum value. After this peak, corresponding to the completion of precipitation, the threshold stress remains constant or even decreases due to limited coarsening of the precipitates. More precise conclusions on this aspect cannot be reasonably formulated at this stage since they require a more extensive microstructural analysis. Nevertheless, recent studies on other materials [26] have indicated that the coarsening process, in the presence of an applied stress, can be governed by complex mechanisms including an acceleration due to the imposition of high stresses.

4.4. The significance of load transfer It is apparent from Figs. 2 and 5 that the threshold stresses in the MMC tend to be larger than in the unreinforced alloy. This may be due to the occurrence of load transfer whereby part of the external load is transferred to the reinforcement and there is a corresponding reduction in the level of the effective stress [27,28]. Recent experiments have shown that load transfer is initiated in the very early stages of deformation prior to the onset of general yielding [29,30] and this suggests the effective stress in the presence of load transfer, |e(LT), may be expressed in the form [5] |e(LT) = (1− h)| − |0,

(9)

where h is the load transfer coefficient having values in the range from 0 in the absence of load transfer to 1 when all of the load is transferred. There may be also an additional strengthening effect due, for example, to an increased dislocation density because of the thermal mismatch between the matrix and the reinforcement. These other factors can be incorporated into the analysis using a relationship analogous to Eq. (9) so that, in practice, h may incorporate other strengthening processes [31,32]. Substitution of Eq. (9) into Eq. (1) leads to the relationship



DGb (1−h)|− |0 m; m = A kT G





n

,



(10)

which, for convenience, may be expressed as DGb(1− h)n |− |*0 m; m = A kT G

n

,

(11)

where |0*= |0/(1− h) represents an apparent threshold stress [5]. It follows from a comparison of Eqs. (1) and (11) that, when the threshold stress is calculated by conventional methods such as by plotting m; 1/n against | on linear axes [33] or by extrapolating the data to a very

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Fig. 9. Temperature-compensated minimum creep rate vs normalized effective stress for an Al-2124 PM + 20% SiC composite [36]: the solid line is reproduced from Fig. 8.

low strain rate [10], these values are more correctly defined as the apparent threshold stress, |0*. A similar argument applies also to the present investigation where the estimated values of |0 in the MMC are more correctly denoted as values of |0*. Since the value of A in Eq. (1) is assumed to be the same for both the unreinforced and the reinforced materials, it follows that, under conditions of constant (| −|0*), the magnitude of the load transfer coefficient, h, may be estimated from the expression [8] m; composite = (1− h)n. m; alloy

(12)

However, it is apparent from Fig. 6 that the unreinforced matrix alloy and the composite exhibit similar creep rates when plotted in terms of the effective stress so that h =0 in the Al-2024 composite and there is no evidence for any load transfer. The results obtained in this investigation for the PM Al-2024 alloy and the Al-2024 composite are therefore very similar to data reported earlier for an unreinforced PM Al-2124 alloy [4] and an Al-2124 alloy reinforced with 10 vol.% SiC particulates [34] where an analysis demonstrated both sets of creep rates superimpose when they are plotted against the effective stress although as in the present investigation, the value of |0 was larger in the composite [31]. For the PM Al-2124 system, the creep rates of the unreinforced alloy and the MMC essentially coincided at very high strain rates in the vicinity of  10 − 1 s − 1 due, it was suggested, to debonding between the SiC particulates and the matrix at these rapid rates [35]. There may be a similar effect in the PM Al-2024 system used in this investigation but, as is evident from inspection of Fig. 2, the tests in both the composite and the unreinforced alloy were not extended to rates faster than 10 − 4 s − 1. Fig. 9 demonstrates a similarity in creep behaviour between the PM Al-2024 materials used in this investi-

gation and a PM Al-2124 alloy reinforced with 20 vol.% SiC particulates [36]. The data are plotted in the form of the normalized minimum creep rate vs the normalized effective stress, where the threshold stresses were recalculated for each temperature by taking n= 4.4 instead of n= 5 as in the original analysis. It is important to note that the solid line in Fig. 9 is not a best fit to the datum points for the Al-2124 composite but rather it is the same line giving a best fit to the experimental points for the PM Al-2024 materials in Fig. 8. It is evident from Fig. 9 that creep data for the Al-2124 and Al-2024 materials essentially superimpose although the creep rates in the Al-2124 composite may be slightly faster by a factor of 2.

5. Summary and conclusions (1) The creep properties of an Al-2024 composite reinforced with 15% SiC particulates were investigated at temperatures from 548 to 603 K. The creep curves for each testing condition show the occurrence of a minimum creep rate prior to an extensive tertiary stage. (2) An analysis of the creep data reveals the presence of a threshold stress. Because of the occurrence of a continuous precipitation of fine particles during each test, the magnitude of this threshold stress depends upon both the testing temperature and the level of the applied stress. (3) When the threshold stress is incorporated into the analysis, it is shown that the stress exponent is 4.4 and the activation energy is close to the value anticipated for lattice self-diffusion in pure aluminium. These results suggest that dislocation climb in the matrix alloy is the rate-controlling creep process. (4) The results for the composite are consistent with those reported earlier for an unreinforced Al-2024 alloy and they are similar also to data reported for a composite having an Al-2124 matrix.

Acknowledgements The research was supported by CNR-Progetto Finalizzato Materiali Speciali per Tecnologie Avanzate II and by the Italian Ministry of University for Scientific and Technological Research.

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