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Liquid impact erosion of Al-Mg and Al-Cu alloys
Wear, 60 (1980) 285 - 304 0 Elsevier Sequoia S.A., Lausanne - Printed in the Netherlands
285
LIQUID IMPACT EROSION OF Al-Mg AND Al-Cu ALLOYS
CAROLYN
M. PREECE
Bell Laboratories,
Murray Hill, N.J. 079 74 (U.S.A.)
JOHN H. BRUNTON Engineering
Department,
Cam bridge University,
Cam bridge (G t. Britain)
RICHARD F. LAWN G.K.N.
Contractors
Ltd., Redditch
(Gt. Britain)
(Received April 26, 1979)
Summary The response of Al-Mg and Al-Cu alloys to multiple liquid impact in the velocity range 90 - 148 m s-l has been investigated. The variation in erosion rate with Al-Mg composition is discussed in terms of a change in the erosion mechanism from ductile rupture of small particles to the removal of whole grains by intercrystalline fracture. This transition occurs with both increased magnesium content and increased velocity of impact. In contrast, Al- 4%~ fails by a mixture of ductile rupture and transgranular fracture The latter mode exhibits river patterns typical of cleavage but occasionally fatigue-like striations are also visible.
1. Introduction Many investigations over the past few decades have shown that there is no general correlation between resistance to liquid impact and any standard mechanical property [ 11. This is not really surprising in view of the fact that mechanical tests are normally carried out at low strain rates and the material is stressed relatively uniformly throughout its bulk, whereas in liquid impact the loading rate is very high and the load is applied to the surface on a local scale. Despite these obvious factors there have been few attempts to relate the erosion rate data to the specific failure modes of the material and in turn to its microstructure. Thus the purpose of the present investigation was to determine the influence of the composition and heat treatment of binary aluminum alloys on the mechanism of material removal by repeated liquid impact and to relate any differences in failure modes to differences in erosion rates.
286 TABLE 1 Compositions and heat treatments of alloys Al-Mg
AI-4%Cu
altoys
Wt.% Mg
At.%
0.45 1.02 2.00 3.00 4.50 9.00
0.54 1.13 2.14 3.40 5.10 10.00
alloys
Mg 4.0 wt.% Cu, 1.74 at.% Cu
Solution-treated
samples
Peak-aged samples
Overaged samples
500 “C for 2 h and ice
water quenched
SAMPLE
540 “C! for 2.5 h and ice water quenched 540 “C for 2.5 h, quenched, aged at 220 “C for 3.25 h 540 “C for 2.5 h, quenched, aged at 425”Gfor 15 h
DRIVE SHAFT TO MOTOR
I
PAOTECllVE
TANK
WHEEL AND JET APPARATUS FOR EROSW STWES (a)
WATER JET (b)
Fig. 1. (a) Schematic illustration of the wheel and jet device for multiple liquid impact. (b) Sample holder for sheet samples for the wheel and jet device.
2. Experimental techniques Disc samples of Al-Mg and Al-& alloys of the compositions and heat treatments given in Table 1 were polished prior to heat treatment and exposed to multiple liquid impact in the system illustrated schematically in Fig. 1. In this system the samples were held in holders which were clamped onto the periphery of a vertical metal wheel in diametrically opposite positions as shown. Two or four samples could be tested simultaneously. A horizontal jet of tap water of diameter 1.5 mm at a head pressure of 20 lbf in-’ was positioned with respect to the wheel so that the jet struck the center of each sample on each revolution of the wheel. Motor speeds, determined by a phototransistor technique, of 95, 125 and 155 Hz corresponding to sample
287
velocities of 91, 119 and 148 m s-l were employed. The system was originally designed for rod samples and the sample holders had to be modified as shown in Fig. l(b) for the sheet samples used in these experiments. This design is not ideal and leads to complex flow patterns. Thus although we consider it’to be satisfactory for determining the relative erosion rates of different materials and identifying the erosion mechanisms, the data may not be comparable with that using other systems. Samples were always exposed to the liquid impact for the same time period (30 s) to ensure that the effects of starting and stopping the water jet were the same for each sample. The damage produced in the samples by the wheel and jet method was studied by weight loss measurements and scanning electron microscopy. Some of the fracture facets produced in Al-Cu and described later were examined by a micro-Laue X-ray technique* to determine whether or not they were crystallographic. In addition, some samples were exposed to single drop impacts at supersonic velocities in the system illustrated in Fig. 2. A small volume of water was contained in a stainless steel chamber. The chamber converged at one end to an orifice of diameter 1.5 mm and the other end, which was open during filling, was sealed by a neoprene disc. The water was extruded through the orifice as a cylindrical column by firing a flat-nosed bullet into a sealed end. The bore of the chamber was 5.3 mm and its interior was polished; changes in cross section were well rounded in order to minimize turbulence [ 21. Measurements of the velocity and behavior of the water drop were obtained using a Cranz-Schardin spark camera [ 31 from which six photographs were obtained at intervals of 13 - 16 ps.
Fig. 2. Schematic
illustration
of the single jet impact
device.
3. Results 3.1. Wheel and jet experiments 3.1.1. Al-Mg alloys Plots of the mass loss of Al-Mg alloys uersus the number of impacts at the three different velocities are shown in Figs. 3 - 5. Each plot is the mean *Courtesy
of the Central
Electricity
Research
Laboratory.
288
16
0 0
IO
20
30
40
50
NUMBER
60 OF
Fig. 3. The mass loss of Al-Mg content: q,0.45%;0, l.O5%;A,
NUMBER
OF
70 IMPACTS
80
90
100
110
XIO-3
alloys subjected to impact 2.00%;m, 3.00%;n,4.50%;@,
IMPACTS
1
I
80
100
at 91 m s-l. 9.00%.
Magnesium
XlO-3
Fig. 4. The mass loss of Al--Mg alloys subjected to impact content: 0,0.45%; n, 3.OO%;A, 4.5O%;e, 9.00%.
at 119 m s--l. Magnesium
of three tests and typical scatter limits are shown by the vertical lines. These figures show that increasing the magnesium content produces a steady increase in the number of impacts required for the onset of reproducibly detectable mass loss but that there is little variation in the maximum mass loss rates with composition. Moreover, at impacts of 91 m s-l, whereas erosion of the alloys containing 3.0% magnesium or less apparently saturates, the Al- 4.5%Mg and Al-9%Mg continue to lose material at a fairly constant or increasing rate.
289
Fig. 5. The mass loss content:
US0.45%;e,
of Al-Mg
alloys subjected to impact at 148 m s-l. Magnesium
3,O%;A, 4.5%;r,
s.o%.
Fig. 6. The development of erosion damage (91 m s-l) in Al-0.45%h!g: (a) 1900 impacts, (b) 2975 impwts” (c) !?85@ impact;;;(d) 3325 ilnpacts, (e) 3809 i:rqacts; (f) 4275 in1pocts.
(a)
(b)
(d)
(e)
(f)
Fig. 7. The development of erosion damage (91 m s-l) in Al-4.5%Mg: (a) 5700 impacts; (b) 11400 impacts;(c) 22 800 impacts;(d) 34 200 impacts;(e) 39 900 impacts;(f) 42 750 impacts.
The reason for this behavior is apparent from the scanning electron micrographs in Figs. 6 - 8. These show the development of erosion damage with increasing numbers of jet impacts for Al-0.45%Mg, Al-4.5%Mg and Al-9%Mg respectively. The initial damage to each of these alloys is in the form of small (about 15 pm) depressions over the whole area of impact. As the number of impacts increases the surface of the Al-0.45%Mg alloy becomes progressively more “undulated” and larger pits develop apparently randomly from some of the small depressions. These pits develop into large (about 50 pm) craters with raised rims (Fig. 6(e)) and erosion occurs by ductile removal of the material constituting the rim (Fig. 6(f)). As erosion continues the damaged area does not grow appreciably but the depth of penetration of the water increases until the sample is actually perforated. The perforation then accounts for the leveling off of the curves in Fig. 3. Micrographs of perforated samples of Al-0.45%Mg and Al-2%Mg are given in Fig. 9. A similar process occurs in the Al- 4_5%Mg but on a finer scale and after a greater number of impacts. Moreover, erosion does not continue by tunneling through the thickness of the sample. Instead the edges of the
(a)
(b)
(d)
(e)
(f)
Fig. 8. The development of erosion damage (91 m s-l) in AliS%Mg: (a) 11400 impacts; (b) 28 500 impacts; (c) 39 900 impacts;(d) 51300 impacts; (e) 57 000 impacts; (f) 62 700 impacts.
eroded area are peeled back as illustrated in Fig. 9. This process causes damage to a larger exposed area than in alloys containing less solute and explains the continued increase in erosion rate of Al- 4.5%Mg shown in Fig. 3. The peeling back also exposes areas which resemble cleavage facets at low magnification but which in fact have a dimpled structure (Fig. 9). The Al-9%Mg behaves quite differently from the other alloys as illustrated in Fig. 8. The initial depressions grow only slowly and do not develop into craters. Instead, material is lost by removal of whole grains by intercrystalline fracture (Fig. 8(e)). At low magnifications the grain boundary surface sometimes (but not always) appears striated (Fig. 8(f)) but at high magnifications the surface is found to be dimpled as shown in Fig. 9. With continued jet impact the eroded area grows laterally by the removal of more surface grains and, like the Al-4.5%Mg samples, the erosion rate continues to increase (Fig. 6). Increasing the velocity of impact from 91 to 119 m s-l has no influence on the mechanism of erosion in alloys containing 0 - 3% magnesium but results in a very small amount of intercrystalline fracture in the Al4_5%Mg samples and less roughening of the surface of the Al-9%Mg prior
292
Al-O.45
% Mg
(a)
Al-2
0%
(b)
Mg
Al-4.5%Mg
(h)
(il
AL - 9.0 % Mg
Fig. 9. The surfaces of Al-0.45%Mg, Al-2.0%Mg, Al-4.5%Mg and Al-9%Mg after a mass loss of about 12 mg. Note that the water has tunneled through the thickness of the Al-0.45%Mg and Al-2.0%Mg alloys whereas the damage spreads over a larger surface area in the more highly alloyed samples. In Al-9%Mg the erosion is intercrystalline.
to fracture. At an impact velocity of 148 m s-l more intercrystalline fracture was observed in the Al- 4.5%Mg alloy and the Al-9%Mg was almost entirely intercrystalline (Fig. 10). 3.1.2. Al-4%Cu Figure 11 shows the mass loss curves for 91 m s-l impacts of Al47%~ in the solution treated and quenched state after peak aging and after
(a)
(b) Fig. 10. The erosion of (a) Al-4.5%Mg
NUMBER
OF IMPACTS
and (b) Al-S.O%Mg by liquid impact at 148 m s-l.
X10-’
Fig. 11. The mass loss of Al-4%Cu subjected to impact at 91 m s-l after different heat treatments: A, solution treated; 0, peak aged; 0, overaged.
12
NUMBER
OF
IMPACTS
XlO-3
Fig. 12. The mass loss of peak aged AI-4%Cu ms -1 .
(a)
(b)
(d)
(e)
subjected
to impact
at 91, 119 and 148
(f)
Fig. 13. The development of erosion damage in solution-treated Al-4%Cu at 91 m SK' ; (a) 42 750 impacts;(b) 51 300 impacts;(c) 59 860 impacts; (d) 68400 impacts; (e) 85 500 impacts; (f) 94 000 impacts.
overaging. It is obvious that heat treatment has little influence on the response of this alloy to jet impact. Because of the similarity in behavior only the peak-aged alloy was tested at higher velocities. The data for this alloy impacted at 91, 119 and 148 m s-l are given in Fig. 12.
295
lopm l----l
Fig. 14. Different
features
observed
in Al-4%Cu
after being impacted
at 91 m s-l.
296
For all heat treatments and at all velocities Al-4%Cu failed by a mixture of ductile erosion of small particles (similar to the low solute content Al-Mg alloys) and transgranular crack propagation, a typical example of which is shown in Fig. 13. The fracture surfaces exhibit several different features illustrated in Fig. 14. These include dimpled structures, finely striated areas and cracks which appear as if they might be intercrystalline but in fact develop in a transcrystalline mode. Another feature observed in Al-4%Cu, but only after impact at 148 m s-r, was the formation of narrow (about 20 pm) tunnels such as those shown in Fig. 15.
Fig. 15. The formation
of small tunnels
in Al-4%Cu
by liquid impact
at 148 m s-l.
The orientation of the facets was determined by a micro-Laue X-ray technique. A typical Laue back-reflection photograph of a “cleavage” facet is shown in Fig. 16(a) and a similar photograph of a region of ductile erosion is given in Fig. 16(b). It is obvious from the former that the film is of a single grain and the sharpness of the spots indicates little deformation of the surface. Figure 16(b), however, indicates a very high degree of deformation so that although the beam (50 pm in diameter) could not have exposed more than about three grains complete Debye rings are obtained. The X-ray photographs of all the facets examined showed the fracture surface to be within 15” of {loo}. When the difficulty of aligning the small (less than 1 mm) facets so that they were exposed by and normal to the X-ray beam is taken into account, it may be reasonably assumed that the facets were actually of a { 100) orientation. 3.2. Single impact tests A typical sequence of photographs of the water jet ejected from the nozzle onto the surface of the sample at about 550 m s-l is given in Fig. 17. For all alloy compositions and heat treatments a single impact produced a rough depression on the surface within which the grain boundaries were made visible by differential deformation in adjacent grains. Subsequent
297
the ~~p~s~o~s uxeil afteraboat fxxk impacts the smple was perforated. This perforation pwcess was entirely ductile for dxfxthe alloys and there was no evidence of &her intercrystalline fracturq as observed fn Al--9%Mg imthe wheel and jet experiment, or the transcrystailine fraet@re exhibit& by Al-4F”Hg in the wheel and jet e~e~me~~. ~~~~~g electron rn~~~~hs of a -&&a3 sampIe, $2 &is case A~-~~~~g~ after teX% impacts are shown in Fig. f83,
impacta~~~e~ed
298
(b)
(a)
(d)
(e) Fig. 17. Movement of a water jet produced in the system illustrated in Fig. 2. The water chamber is on the right and the sample on the left. The times between frames (a), (b), (c) and (d) are 15, 13.7 and 13.2 /JS respectively.
p=plclv
i
p2c2 P,C,
+pzcz
(1) 1
where p, c and u represent the density, velocity of sound and velocity of impact respectively and the subscripts 1 and 2 pertain to water and the metal. This relation is not accurate for the present study because it assumes (a) a flat slug of liquid and (b) only an elastic response from the target, whereas plastic deformation actually occurs. This equation gives the maximum pressure occurring at the center line of impact, whereas for a spherical or cylindrical drop the pressure rises on either side of the line, as the contact area increases, to a value in the region of 1.8P [8] to 3P [9] just prior to the lateral flow of water outwards. The second cause of damage is the flow of water across the metal surface after impact. Rochester and Brunton [8] have shown that the magni-
299
(c)
20pm
\
-
(e)
\
Fig. 18. The surface of Al-Q%Mg after ten impacts at about 550 m sdi. Note the ductile perforation of the sample and the absence of any intercrystalline fracture.
tude of the shear stress imposed by the flow over a smooth surface is insignificant compared with the impact pressures. However, a roughened surface suffers significant damage by this process; the stress produced by the outward flow on a surface step or protuberance exceeds the impact pressure
300
by a significant amount because of the increase in the velocity of the liquid on jetting. In the present experiments the pressure given by eqn. (1) corresponds to 130,170,211 and 786 MPa for impact velocities of 91, 119, 148 and 550 m s-l respectively. The quasi-static tensile yield and fracture stresses of Al-Mg and Al-Cu alloys are given in Table 2 for comparison purposes. Even at the lowest velocity the water hammer pressure is greater than the yield stress of all the alloys except the peak-aged Al-4%Cu alloy. If a maximum pressure of 1.W is attained, then this is greater than the fracture stress of all the alloys except the Al-9%Mg. The fact that the load is applied as a compression rather than as a tension probably accounts for the observation that many impacts are required before any mass loss can be detected. TABLE
2
Quasi-static
tensile
yield and fracture
stresses of Al-Mg
and Al-h
alloys
Composition
Yield stress W’a)
Fracture stress (UTS) WPa)
Al-0.45%Mg Al-1.02%Mg Al-Z.O%Mg Al-3.0%Mg Al-4.5%Mg Al-S.O%Mg Al-4.O%Cu, treated Al-4.0%Cu,
solution
17.25 21.5 34.5 53.0 73.5 120.5 119.0
79.5 92.7 138.0 174.4 224 317 213
peak aged
149.0
238
The initial response of all the alloys, i.e. a shallow pitting of the surface, is typical of metals exposed to liquid impact. Thomas [lo] has shown that these are not due to spikes on the compression wave produced for example by cavitation in the liquid. Thus he attributed the pitting to inhomogeneities in the strength of the metal surface. As also suggested by Thomas [lo] it should be noted that unlike most loading conditions, e.g. impact by a rigid indenter, in which the load is borne mostly by the strongest parts of the loaded area the liquid impact load is applied equally over the whole area of impact and the resistance to deformation is determined by the weakest areas. In other words the weak areas are not prevented from deforming by for example the contact of stronger areas with a rigid indenter. Once depressions have formed - in a weak phase or in the vicinity of favorably oriented dislocation sources - they have two consequences in the low strength alloys. Firstly the impinging water is more effectively trapped in the depressions than elsewhere since outward flow is limited by the walls of the depression. This increases the duration of the load in these areas and results in the “tunneling” effect observed in the softer samples (Fig. 9). Secondly the depressions give rise to a roughened surface and any raised rims
301
or other protuberances produced by the formation of depressions will be subjected to severe shear forces from the outward flow of water across the surface. In the higher strength aluminum alloys the uniform mode of loading manifests itself in different fracture modes. Both Al-9%Mg and Al-4%Cu exhibit crack initiation and growth rather than simply the ductile shearing off of small particles. However, the cracks are along or adjacent to the grain boundaries in the former alloy and are transgranular and apparently crystallographic in the latter. This difference is most probably attributable to the fact that Al-Mg alloys do not form a series of pre-precipitates as do Al-Cu and many other aluminum alloys. The only intermediate precipitate which has been observed in Al-Mg, i.e. fl’, has a larger specific volume than the matrix. It is therefore unable to precipitate in the vacancydeficient regions adjacent to the grain boundaries, leaving a large precipitate-free zone (PFZ) [ 111. Moreover, nucleation of both p’ and the equilibrium p phase Mg, Ala is difficult and results in a very coarse precipate distribution compared with that occurring in other aluminum alloys [ 121. Although the AI-Mg alloys in this experiment were solution treated and quenched, precipitation could not be inhibited in the samples containing 9% magnesium. Optical metallography showed these samples to have a coarse microstructure with preferential grain boundary precipitation and evidence of a PFZ. This zone is obviously the weakest part of the structure and the region where deformation and crack growth can be facilitated. The striated and dimpled nature of the fracture surface in this alloy indicates an intermittent crack propagation with considerable ductile deformation ahead of the crack. Similar intergranular failure has been observed in aged Al-7%Mg during tensile loading [12] and the loading conditions under liquid impact maximize this effect. The tap water in Cambridge is particularly “hard” and there was some suspicion that stress corrosion cracking might be the cause of the intercrystalline failure mode. Therefore a few tests were performed using deionized water. Since a maximum head pressure of only 5 lbf inM2 was obtainable for the tank containing the deionized water, the experiments were also repeated using the same tank and tap water. The effect of the lower head pressure was to decrease the rate of erosion, but the fracture surfaces were indistinguishable from each other (Fig. 19) and from those obtained with tap water at 20 lbf ine2 (Fig. 10). In contrast to the Al-Mg, a sequence of Guinier-Preston zones, coherent, semi-coherent and incoherent precipitates occurs in the Cu-Al system leading to a fine dispersion of the second phase. Moreover, the nucleation and growth of 8 ’ in Al-4%Cu requires the emission of vacancies, and this is obviously an easier process in the vacancy-deficient areas near the grain boundaries than in the vacancy-saturated interior of the grain. Thus in contrast to the 8’ phase in Al-Mg the 0 ’ precipitates preferentially near the grain boundaries [ 111 and there is no significant PFZ; hence there is no region which is susceptible to easy deformation as there is in the Al-9%Mg. Moreover, the precipitate structure in all the Al-4%Cu alloys was too fine to be resolved
(a)
(b)
(cl
(d)
Fig. 19. The surfaces of Al-9%Mg eroded head pressure of 5 Ibf ine2 and an impact
in (a) tap water and (b) deionized velocity of 148 m s-l.
water at a
in the optical microscope. The dispersion was apparently also too fine to permit sufficient ductility for the growth and coalescence of voids and dimpled fracture surfaces are rarely observed. The (100) crystallographic nature of the fracture facets in Al-4%Cu suggests either a brittle fracture mode or a fatigue-type failure. Brittle striations have been observed on the fatigue fracture surfaces of several aluminum alloys [13 - 171. While these have been attributed by some to alternating slip on intersecting (111) planes resulting in the ductile propagation of a crack along the { 100) [ 171, others have attributed them to a true cleavage fracture [ 151. Cleavage of aluminum has been observed in liquid metal embrittlement studies [18, 191 and in corrosion fatigue [20], but there is both theoretical and experimental controversy about the specific fracture planes since both {loo} and (111) planes have been both observed [ 15,18,19] and predicted [21, 221. In the present study both ductile and “brittle” failure modes occurred in the same sample and it would therefore
303
seem more likely that the latter is actually pseudo-cleavage resulting from slip on intersecting planes. However, the sharpness of the X-ray spots obtained from the brittle regions tends to dispute this idea. It is therefore obvious that this aspect of the study needs further investigation. An attempt to correlate the striation spacing in either Al-9%Mg or Al-4%Cu with the number of impacts has not been made because it is not known if the crack growth observed after a particular period of exposure to the liquid impact occurred gradually over the whole period, or intermittently or catastrophically after one or a few impacts. The effect of the variation in fracture modes with composition of these alloys is to preclude any simple correlation between incubation period or rate of mass loss and any quasi-static strength parameter. The intercrystalline fracture of Al- 9%Mg and the “brittle” fracture of Al-4%Cu lead to considerably higher rates of mass loss than expected from their strength or hardness. In the high velocity single-impact studies the Al-9%Mg and Al-4%Cu alloys had been expected to fail in a similar manner to that in the wheel and jet experiments. The penetration of samples of each alloy in less than ten impacts, however, obviously precluded any intermittent crack growth such as that responsible for the failures by the wheel and jet. The ductile tearing of the samples at the high velocity must therefore be regarded as the dynamic failure mode of all the alloys.
Acknowledgments The assistance of Dr. E. Metcalfe of the Central Electricity Research Laboratory, Leatherhead, in obtaining the micro-Laue X-ray photographs is greatly appreciated. One of the authors (C.M.P.) is also grateful to the John Simon Guggenheim Foundation and to Churchill College Cambridge for Fellowships during the period of this research. The work was supported by the U.K. Scientific Research Council.
References 1 J. H. Brunton, in C. M. Preece (ed.), Erosion, Treatise on Materials Science and Technology, Vol. 16, Academic Press, New York, 1979, p. 185. 2 J, H. Brunton, Philos. Trans. R. Sot. London, Ser. A, 260 (1977) 79. 3 C. Cranz and H. Schardin, 2. Phys., 56 (1929) 147. 4 W. F. Adler, in C. M. Preece (ed.), Erosion, Treatise on Materials Science and Technology, Vol. 16, Academic Press, New York, 1979, p. 127. 5 P. de Haller, Schweiz. Bauztg., 101 (1933) 243, 260. 6 0. G. Engel,J. Res. Nat. But-. Stand., 54 (1955) 51. 7 F. P. Bowden and J. H. Brunton, Proc. R. Sot. London, Ser. A, 263 (1961) 433. 8 M. C. Rochester and J. H. Brunton, in A. A. Fyall and R. B. King (eds.), Proc. 4th Int. Conf on Rain Erosion and Allied Phenomena, Royal Aircraft Establishment, Farnborough, 1974, p. 371.
304 9 10 11 12 13 14 15 16 17 18 19 20 21 22
F. J. Heymann, J. Appi. Phys., 40 (1969) 5113. G. P. Thomas, Philos. Trans. R. Sot. London, Ser. A, 260 (1966) 140. A. Kelly and R. B. Nicholson, Prog. Mater. Sci., 10 (3) (1963) 245. G. Thomas and J. Nutting, J. Inst. Mef., 86 (1957 - 58) 7. P. J. E. Forsyth, C. A. Stubbington and D. Clark, J. Inst. Met., 90 (1961 - 2) 238. P. J. E. Forsyth, J. Inst. Met., 93 (1965) 456. K. R. L. Thomspon and J. V. Craig, Metall. Trans., 1 (1970) 1047. C. Laird,Am. Sot. Test. Muter. Spec. Tech. Publ., 415, 1967. G. G. Garrett and J. F. Knott, Act4 Met&l., 23 (1975) 841. A. R. C. Westwood, C. M. Preece and 111.H. Kamdar, Trans. Am. Sot. Met., 60 (1967) 263. R. G. Davies and N. S. Stoloff, Trans. MetalI. Sot. AZME, 233 (1965) 714. C. A. Stubbington, Metullurgiu, 68 (1963) 109. J. J. Gilman, in Fracture, Wiley, New York, 1959, p. 193. J. J. Burton and G. Jura, J. Phys. Chem., 17 (1967) 1937.
285
LIQUID IMPACT EROSION OF Al-Mg AND Al-Cu ALLOYS
CAROLYN
M. PREECE
Bell Laboratories,
Murray Hill, N.J. 079 74 (U.S.A.)
JOHN H. BRUNTON Engineering
Department,
Cam bridge University,
Cam bridge (G t. Britain)
RICHARD F. LAWN G.K.N.
Contractors
Ltd., Redditch
(Gt. Britain)
(Received April 26, 1979)
Summary The response of Al-Mg and Al-Cu alloys to multiple liquid impact in the velocity range 90 - 148 m s-l has been investigated. The variation in erosion rate with Al-Mg composition is discussed in terms of a change in the erosion mechanism from ductile rupture of small particles to the removal of whole grains by intercrystalline fracture. This transition occurs with both increased magnesium content and increased velocity of impact. In contrast, Al- 4%~ fails by a mixture of ductile rupture and transgranular fracture The latter mode exhibits river patterns typical of cleavage but occasionally fatigue-like striations are also visible.
1. Introduction Many investigations over the past few decades have shown that there is no general correlation between resistance to liquid impact and any standard mechanical property [ 11. This is not really surprising in view of the fact that mechanical tests are normally carried out at low strain rates and the material is stressed relatively uniformly throughout its bulk, whereas in liquid impact the loading rate is very high and the load is applied to the surface on a local scale. Despite these obvious factors there have been few attempts to relate the erosion rate data to the specific failure modes of the material and in turn to its microstructure. Thus the purpose of the present investigation was to determine the influence of the composition and heat treatment of binary aluminum alloys on the mechanism of material removal by repeated liquid impact and to relate any differences in failure modes to differences in erosion rates.
286 TABLE 1 Compositions and heat treatments of alloys Al-Mg
AI-4%Cu
altoys
Wt.% Mg
At.%
0.45 1.02 2.00 3.00 4.50 9.00
0.54 1.13 2.14 3.40 5.10 10.00
alloys
Mg 4.0 wt.% Cu, 1.74 at.% Cu
Solution-treated
samples
Peak-aged samples
Overaged samples
500 “C for 2 h and ice
water quenched
SAMPLE
540 “C! for 2.5 h and ice water quenched 540 “C for 2.5 h, quenched, aged at 220 “C for 3.25 h 540 “C for 2.5 h, quenched, aged at 425”Gfor 15 h
DRIVE SHAFT TO MOTOR
I
PAOTECllVE
TANK
WHEEL AND JET APPARATUS FOR EROSW STWES (a)
WATER JET (b)
Fig. 1. (a) Schematic illustration of the wheel and jet device for multiple liquid impact. (b) Sample holder for sheet samples for the wheel and jet device.
2. Experimental techniques Disc samples of Al-Mg and Al-& alloys of the compositions and heat treatments given in Table 1 were polished prior to heat treatment and exposed to multiple liquid impact in the system illustrated schematically in Fig. 1. In this system the samples were held in holders which were clamped onto the periphery of a vertical metal wheel in diametrically opposite positions as shown. Two or four samples could be tested simultaneously. A horizontal jet of tap water of diameter 1.5 mm at a head pressure of 20 lbf in-’ was positioned with respect to the wheel so that the jet struck the center of each sample on each revolution of the wheel. Motor speeds, determined by a phototransistor technique, of 95, 125 and 155 Hz corresponding to sample
287
velocities of 91, 119 and 148 m s-l were employed. The system was originally designed for rod samples and the sample holders had to be modified as shown in Fig. l(b) for the sheet samples used in these experiments. This design is not ideal and leads to complex flow patterns. Thus although we consider it’to be satisfactory for determining the relative erosion rates of different materials and identifying the erosion mechanisms, the data may not be comparable with that using other systems. Samples were always exposed to the liquid impact for the same time period (30 s) to ensure that the effects of starting and stopping the water jet were the same for each sample. The damage produced in the samples by the wheel and jet method was studied by weight loss measurements and scanning electron microscopy. Some of the fracture facets produced in Al-Cu and described later were examined by a micro-Laue X-ray technique* to determine whether or not they were crystallographic. In addition, some samples were exposed to single drop impacts at supersonic velocities in the system illustrated in Fig. 2. A small volume of water was contained in a stainless steel chamber. The chamber converged at one end to an orifice of diameter 1.5 mm and the other end, which was open during filling, was sealed by a neoprene disc. The water was extruded through the orifice as a cylindrical column by firing a flat-nosed bullet into a sealed end. The bore of the chamber was 5.3 mm and its interior was polished; changes in cross section were well rounded in order to minimize turbulence [ 21. Measurements of the velocity and behavior of the water drop were obtained using a Cranz-Schardin spark camera [ 31 from which six photographs were obtained at intervals of 13 - 16 ps.
Fig. 2. Schematic
illustration
of the single jet impact
device.
3. Results 3.1. Wheel and jet experiments 3.1.1. Al-Mg alloys Plots of the mass loss of Al-Mg alloys uersus the number of impacts at the three different velocities are shown in Figs. 3 - 5. Each plot is the mean *Courtesy
of the Central
Electricity
Research
Laboratory.
288
16
0 0
IO
20
30
40
50
NUMBER
60 OF
Fig. 3. The mass loss of Al-Mg content: q,0.45%;0, l.O5%;A,
NUMBER
OF
70 IMPACTS
80
90
100
110
XIO-3
alloys subjected to impact 2.00%;m, 3.00%;n,4.50%;@,
IMPACTS
1
I
80
100
at 91 m s-l. 9.00%.
Magnesium
XlO-3
Fig. 4. The mass loss of Al--Mg alloys subjected to impact content: 0,0.45%; n, 3.OO%;A, 4.5O%;e, 9.00%.
at 119 m s--l. Magnesium
of three tests and typical scatter limits are shown by the vertical lines. These figures show that increasing the magnesium content produces a steady increase in the number of impacts required for the onset of reproducibly detectable mass loss but that there is little variation in the maximum mass loss rates with composition. Moreover, at impacts of 91 m s-l, whereas erosion of the alloys containing 3.0% magnesium or less apparently saturates, the Al- 4.5%Mg and Al-9%Mg continue to lose material at a fairly constant or increasing rate.
289
Fig. 5. The mass loss content:
US0.45%;e,
of Al-Mg
alloys subjected to impact at 148 m s-l. Magnesium
3,O%;A, 4.5%;r,
s.o%.
Fig. 6. The development of erosion damage (91 m s-l) in Al-0.45%h!g: (a) 1900 impacts, (b) 2975 impwts” (c) !?85@ impact;;;(d) 3325 ilnpacts, (e) 3809 i:rqacts; (f) 4275 in1pocts.
(a)
(b)
(d)
(e)
(f)
Fig. 7. The development of erosion damage (91 m s-l) in Al-4.5%Mg: (a) 5700 impacts; (b) 11400 impacts;(c) 22 800 impacts;(d) 34 200 impacts;(e) 39 900 impacts;(f) 42 750 impacts.
The reason for this behavior is apparent from the scanning electron micrographs in Figs. 6 - 8. These show the development of erosion damage with increasing numbers of jet impacts for Al-0.45%Mg, Al-4.5%Mg and Al-9%Mg respectively. The initial damage to each of these alloys is in the form of small (about 15 pm) depressions over the whole area of impact. As the number of impacts increases the surface of the Al-0.45%Mg alloy becomes progressively more “undulated” and larger pits develop apparently randomly from some of the small depressions. These pits develop into large (about 50 pm) craters with raised rims (Fig. 6(e)) and erosion occurs by ductile removal of the material constituting the rim (Fig. 6(f)). As erosion continues the damaged area does not grow appreciably but the depth of penetration of the water increases until the sample is actually perforated. The perforation then accounts for the leveling off of the curves in Fig. 3. Micrographs of perforated samples of Al-0.45%Mg and Al-2%Mg are given in Fig. 9. A similar process occurs in the Al- 4_5%Mg but on a finer scale and after a greater number of impacts. Moreover, erosion does not continue by tunneling through the thickness of the sample. Instead the edges of the
(a)
(b)
(d)
(e)
(f)
Fig. 8. The development of erosion damage (91 m s-l) in AliS%Mg: (a) 11400 impacts; (b) 28 500 impacts; (c) 39 900 impacts;(d) 51300 impacts; (e) 57 000 impacts; (f) 62 700 impacts.
eroded area are peeled back as illustrated in Fig. 9. This process causes damage to a larger exposed area than in alloys containing less solute and explains the continued increase in erosion rate of Al- 4.5%Mg shown in Fig. 3. The peeling back also exposes areas which resemble cleavage facets at low magnification but which in fact have a dimpled structure (Fig. 9). The Al-9%Mg behaves quite differently from the other alloys as illustrated in Fig. 8. The initial depressions grow only slowly and do not develop into craters. Instead, material is lost by removal of whole grains by intercrystalline fracture (Fig. 8(e)). At low magnifications the grain boundary surface sometimes (but not always) appears striated (Fig. 8(f)) but at high magnifications the surface is found to be dimpled as shown in Fig. 9. With continued jet impact the eroded area grows laterally by the removal of more surface grains and, like the Al-4.5%Mg samples, the erosion rate continues to increase (Fig. 6). Increasing the velocity of impact from 91 to 119 m s-l has no influence on the mechanism of erosion in alloys containing 0 - 3% magnesium but results in a very small amount of intercrystalline fracture in the Al4_5%Mg samples and less roughening of the surface of the Al-9%Mg prior
292
Al-O.45
% Mg
(a)
Al-2
0%
(b)
Mg
Al-4.5%Mg
(h)
(il
AL - 9.0 % Mg
Fig. 9. The surfaces of Al-0.45%Mg, Al-2.0%Mg, Al-4.5%Mg and Al-9%Mg after a mass loss of about 12 mg. Note that the water has tunneled through the thickness of the Al-0.45%Mg and Al-2.0%Mg alloys whereas the damage spreads over a larger surface area in the more highly alloyed samples. In Al-9%Mg the erosion is intercrystalline.
to fracture. At an impact velocity of 148 m s-l more intercrystalline fracture was observed in the Al- 4.5%Mg alloy and the Al-9%Mg was almost entirely intercrystalline (Fig. 10). 3.1.2. Al-4%Cu Figure 11 shows the mass loss curves for 91 m s-l impacts of Al47%~ in the solution treated and quenched state after peak aging and after
(a)
(b) Fig. 10. The erosion of (a) Al-4.5%Mg
NUMBER
OF IMPACTS
and (b) Al-S.O%Mg by liquid impact at 148 m s-l.
X10-’
Fig. 11. The mass loss of Al-4%Cu subjected to impact at 91 m s-l after different heat treatments: A, solution treated; 0, peak aged; 0, overaged.
12
NUMBER
OF
IMPACTS
XlO-3
Fig. 12. The mass loss of peak aged AI-4%Cu ms -1 .
(a)
(b)
(d)
(e)
subjected
to impact
at 91, 119 and 148
(f)
Fig. 13. The development of erosion damage in solution-treated Al-4%Cu at 91 m SK' ; (a) 42 750 impacts;(b) 51 300 impacts;(c) 59 860 impacts; (d) 68400 impacts; (e) 85 500 impacts; (f) 94 000 impacts.
overaging. It is obvious that heat treatment has little influence on the response of this alloy to jet impact. Because of the similarity in behavior only the peak-aged alloy was tested at higher velocities. The data for this alloy impacted at 91, 119 and 148 m s-l are given in Fig. 12.
295
lopm l----l
Fig. 14. Different
features
observed
in Al-4%Cu
after being impacted
at 91 m s-l.
296
For all heat treatments and at all velocities Al-4%Cu failed by a mixture of ductile erosion of small particles (similar to the low solute content Al-Mg alloys) and transgranular crack propagation, a typical example of which is shown in Fig. 13. The fracture surfaces exhibit several different features illustrated in Fig. 14. These include dimpled structures, finely striated areas and cracks which appear as if they might be intercrystalline but in fact develop in a transcrystalline mode. Another feature observed in Al-4%Cu, but only after impact at 148 m s-r, was the formation of narrow (about 20 pm) tunnels such as those shown in Fig. 15.
Fig. 15. The formation
of small tunnels
in Al-4%Cu
by liquid impact
at 148 m s-l.
The orientation of the facets was determined by a micro-Laue X-ray technique. A typical Laue back-reflection photograph of a “cleavage” facet is shown in Fig. 16(a) and a similar photograph of a region of ductile erosion is given in Fig. 16(b). It is obvious from the former that the film is of a single grain and the sharpness of the spots indicates little deformation of the surface. Figure 16(b), however, indicates a very high degree of deformation so that although the beam (50 pm in diameter) could not have exposed more than about three grains complete Debye rings are obtained. The X-ray photographs of all the facets examined showed the fracture surface to be within 15” of {loo}. When the difficulty of aligning the small (less than 1 mm) facets so that they were exposed by and normal to the X-ray beam is taken into account, it may be reasonably assumed that the facets were actually of a { 100) orientation. 3.2. Single impact tests A typical sequence of photographs of the water jet ejected from the nozzle onto the surface of the sample at about 550 m s-l is given in Fig. 17. For all alloy compositions and heat treatments a single impact produced a rough depression on the surface within which the grain boundaries were made visible by differential deformation in adjacent grains. Subsequent
297
the ~~p~s~o~s uxeil afteraboat fxxk impacts the smple was perforated. This perforation pwcess was entirely ductile for dxfxthe alloys and there was no evidence of &her intercrystalline fracturq as observed fn Al--9%Mg imthe wheel and jet experiment, or the transcrystailine fraet@re exhibit& by Al-4F”Hg in the wheel and jet e~e~me~~. ~~~~~g electron rn~~~~hs of a -&&a3 sampIe, $2 &is case A~-~~~~g~ after teX% impacts are shown in Fig. f83,
impacta~~~e~ed
298
(b)
(a)
(d)
(e) Fig. 17. Movement of a water jet produced in the system illustrated in Fig. 2. The water chamber is on the right and the sample on the left. The times between frames (a), (b), (c) and (d) are 15, 13.7 and 13.2 /JS respectively.
p=plclv
i
p2c2 P,C,
+pzcz
(1) 1
where p, c and u represent the density, velocity of sound and velocity of impact respectively and the subscripts 1 and 2 pertain to water and the metal. This relation is not accurate for the present study because it assumes (a) a flat slug of liquid and (b) only an elastic response from the target, whereas plastic deformation actually occurs. This equation gives the maximum pressure occurring at the center line of impact, whereas for a spherical or cylindrical drop the pressure rises on either side of the line, as the contact area increases, to a value in the region of 1.8P [8] to 3P [9] just prior to the lateral flow of water outwards. The second cause of damage is the flow of water across the metal surface after impact. Rochester and Brunton [8] have shown that the magni-
299
(c)
20pm
\
-
(e)
\
Fig. 18. The surface of Al-Q%Mg after ten impacts at about 550 m sdi. Note the ductile perforation of the sample and the absence of any intercrystalline fracture.
tude of the shear stress imposed by the flow over a smooth surface is insignificant compared with the impact pressures. However, a roughened surface suffers significant damage by this process; the stress produced by the outward flow on a surface step or protuberance exceeds the impact pressure
300
by a significant amount because of the increase in the velocity of the liquid on jetting. In the present experiments the pressure given by eqn. (1) corresponds to 130,170,211 and 786 MPa for impact velocities of 91, 119, 148 and 550 m s-l respectively. The quasi-static tensile yield and fracture stresses of Al-Mg and Al-Cu alloys are given in Table 2 for comparison purposes. Even at the lowest velocity the water hammer pressure is greater than the yield stress of all the alloys except the peak-aged Al-4%Cu alloy. If a maximum pressure of 1.W is attained, then this is greater than the fracture stress of all the alloys except the Al-9%Mg. The fact that the load is applied as a compression rather than as a tension probably accounts for the observation that many impacts are required before any mass loss can be detected. TABLE
2
Quasi-static
tensile
yield and fracture
stresses of Al-Mg
and Al-h
alloys
Composition
Yield stress W’a)
Fracture stress (UTS) WPa)
Al-0.45%Mg Al-1.02%Mg Al-Z.O%Mg Al-3.0%Mg Al-4.5%Mg Al-S.O%Mg Al-4.O%Cu, treated Al-4.0%Cu,
solution
17.25 21.5 34.5 53.0 73.5 120.5 119.0
79.5 92.7 138.0 174.4 224 317 213
peak aged
149.0
238
The initial response of all the alloys, i.e. a shallow pitting of the surface, is typical of metals exposed to liquid impact. Thomas [lo] has shown that these are not due to spikes on the compression wave produced for example by cavitation in the liquid. Thus he attributed the pitting to inhomogeneities in the strength of the metal surface. As also suggested by Thomas [lo] it should be noted that unlike most loading conditions, e.g. impact by a rigid indenter, in which the load is borne mostly by the strongest parts of the loaded area the liquid impact load is applied equally over the whole area of impact and the resistance to deformation is determined by the weakest areas. In other words the weak areas are not prevented from deforming by for example the contact of stronger areas with a rigid indenter. Once depressions have formed - in a weak phase or in the vicinity of favorably oriented dislocation sources - they have two consequences in the low strength alloys. Firstly the impinging water is more effectively trapped in the depressions than elsewhere since outward flow is limited by the walls of the depression. This increases the duration of the load in these areas and results in the “tunneling” effect observed in the softer samples (Fig. 9). Secondly the depressions give rise to a roughened surface and any raised rims
301
or other protuberances produced by the formation of depressions will be subjected to severe shear forces from the outward flow of water across the surface. In the higher strength aluminum alloys the uniform mode of loading manifests itself in different fracture modes. Both Al-9%Mg and Al-4%Cu exhibit crack initiation and growth rather than simply the ductile shearing off of small particles. However, the cracks are along or adjacent to the grain boundaries in the former alloy and are transgranular and apparently crystallographic in the latter. This difference is most probably attributable to the fact that Al-Mg alloys do not form a series of pre-precipitates as do Al-Cu and many other aluminum alloys. The only intermediate precipitate which has been observed in Al-Mg, i.e. fl’, has a larger specific volume than the matrix. It is therefore unable to precipitate in the vacancydeficient regions adjacent to the grain boundaries, leaving a large precipitate-free zone (PFZ) [ 111. Moreover, nucleation of both p’ and the equilibrium p phase Mg, Ala is difficult and results in a very coarse precipate distribution compared with that occurring in other aluminum alloys [ 121. Although the AI-Mg alloys in this experiment were solution treated and quenched, precipitation could not be inhibited in the samples containing 9% magnesium. Optical metallography showed these samples to have a coarse microstructure with preferential grain boundary precipitation and evidence of a PFZ. This zone is obviously the weakest part of the structure and the region where deformation and crack growth can be facilitated. The striated and dimpled nature of the fracture surface in this alloy indicates an intermittent crack propagation with considerable ductile deformation ahead of the crack. Similar intergranular failure has been observed in aged Al-7%Mg during tensile loading [12] and the loading conditions under liquid impact maximize this effect. The tap water in Cambridge is particularly “hard” and there was some suspicion that stress corrosion cracking might be the cause of the intercrystalline failure mode. Therefore a few tests were performed using deionized water. Since a maximum head pressure of only 5 lbf inM2 was obtainable for the tank containing the deionized water, the experiments were also repeated using the same tank and tap water. The effect of the lower head pressure was to decrease the rate of erosion, but the fracture surfaces were indistinguishable from each other (Fig. 19) and from those obtained with tap water at 20 lbf ine2 (Fig. 10). In contrast to the Al-Mg, a sequence of Guinier-Preston zones, coherent, semi-coherent and incoherent precipitates occurs in the Cu-Al system leading to a fine dispersion of the second phase. Moreover, the nucleation and growth of 8 ’ in Al-4%Cu requires the emission of vacancies, and this is obviously an easier process in the vacancy-deficient areas near the grain boundaries than in the vacancy-saturated interior of the grain. Thus in contrast to the 8’ phase in Al-Mg the 0 ’ precipitates preferentially near the grain boundaries [ 111 and there is no significant PFZ; hence there is no region which is susceptible to easy deformation as there is in the Al-9%Mg. Moreover, the precipitate structure in all the Al-4%Cu alloys was too fine to be resolved
(a)
(b)
(cl
(d)
Fig. 19. The surfaces of Al-9%Mg eroded head pressure of 5 Ibf ine2 and an impact
in (a) tap water and (b) deionized velocity of 148 m s-l.
water at a
in the optical microscope. The dispersion was apparently also too fine to permit sufficient ductility for the growth and coalescence of voids and dimpled fracture surfaces are rarely observed. The (100) crystallographic nature of the fracture facets in Al-4%Cu suggests either a brittle fracture mode or a fatigue-type failure. Brittle striations have been observed on the fatigue fracture surfaces of several aluminum alloys [13 - 171. While these have been attributed by some to alternating slip on intersecting (111) planes resulting in the ductile propagation of a crack along the { 100) [ 171, others have attributed them to a true cleavage fracture [ 151. Cleavage of aluminum has been observed in liquid metal embrittlement studies [18, 191 and in corrosion fatigue [20], but there is both theoretical and experimental controversy about the specific fracture planes since both {loo} and (111) planes have been both observed [ 15,18,19] and predicted [21, 221. In the present study both ductile and “brittle” failure modes occurred in the same sample and it would therefore
303
seem more likely that the latter is actually pseudo-cleavage resulting from slip on intersecting planes. However, the sharpness of the X-ray spots obtained from the brittle regions tends to dispute this idea. It is therefore obvious that this aspect of the study needs further investigation. An attempt to correlate the striation spacing in either Al-9%Mg or Al-4%Cu with the number of impacts has not been made because it is not known if the crack growth observed after a particular period of exposure to the liquid impact occurred gradually over the whole period, or intermittently or catastrophically after one or a few impacts. The effect of the variation in fracture modes with composition of these alloys is to preclude any simple correlation between incubation period or rate of mass loss and any quasi-static strength parameter. The intercrystalline fracture of Al- 9%Mg and the “brittle” fracture of Al-4%Cu lead to considerably higher rates of mass loss than expected from their strength or hardness. In the high velocity single-impact studies the Al-9%Mg and Al-4%Cu alloys had been expected to fail in a similar manner to that in the wheel and jet experiments. The penetration of samples of each alloy in less than ten impacts, however, obviously precluded any intermittent crack growth such as that responsible for the failures by the wheel and jet. The ductile tearing of the samples at the high velocity must therefore be regarded as the dynamic failure mode of all the alloys.
Acknowledgments The assistance of Dr. E. Metcalfe of the Central Electricity Research Laboratory, Leatherhead, in obtaining the micro-Laue X-ray photographs is greatly appreciated. One of the authors (C.M.P.) is also grateful to the John Simon Guggenheim Foundation and to Churchill College Cambridge for Fellowships during the period of this research. The work was supported by the U.K. Scientific Research Council.
References 1 J. H. Brunton, in C. M. Preece (ed.), Erosion, Treatise on Materials Science and Technology, Vol. 16, Academic Press, New York, 1979, p. 185. 2 J, H. Brunton, Philos. Trans. R. Sot. London, Ser. A, 260 (1977) 79. 3 C. Cranz and H. Schardin, 2. Phys., 56 (1929) 147. 4 W. F. Adler, in C. M. Preece (ed.), Erosion, Treatise on Materials Science and Technology, Vol. 16, Academic Press, New York, 1979, p. 127. 5 P. de Haller, Schweiz. Bauztg., 101 (1933) 243, 260. 6 0. G. Engel,J. Res. Nat. But-. Stand., 54 (1955) 51. 7 F. P. Bowden and J. H. Brunton, Proc. R. Sot. London, Ser. A, 263 (1961) 433. 8 M. C. Rochester and J. H. Brunton, in A. A. Fyall and R. B. King (eds.), Proc. 4th Int. Conf on Rain Erosion and Allied Phenomena, Royal Aircraft Establishment, Farnborough, 1974, p. 371.
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F. J. Heymann, J. Appi. Phys., 40 (1969) 5113. G. P. Thomas, Philos. Trans. R. Sot. London, Ser. A, 260 (1966) 140. A. Kelly and R. B. Nicholson, Prog. Mater. Sci., 10 (3) (1963) 245. G. Thomas and J. Nutting, J. Inst. Mef., 86 (1957 - 58) 7. P. J. E. Forsyth, C. A. Stubbington and D. Clark, J. Inst. Met., 90 (1961 - 2) 238. P. J. E. Forsyth, J. Inst. Met., 93 (1965) 456. K. R. L. Thomspon and J. V. Craig, Metall. Trans., 1 (1970) 1047. C. Laird,Am. Sot. Test. Muter. Spec. Tech. Publ., 415, 1967. G. G. Garrett and J. F. Knott, Act4 Met&l., 23 (1975) 841. A. R. C. Westwood, C. M. Preece and 111.H. Kamdar, Trans. Am. Sot. Met., 60 (1967) 263. R. G. Davies and N. S. Stoloff, Trans. MetalI. Sot. AZME, 233 (1965) 714. C. A. Stubbington, Metullurgiu, 68 (1963) 109. J. J. Gilman, in Fracture, Wiley, New York, 1959, p. 193. J. J. Burton and G. Jura, J. Phys. Chem., 17 (1967) 1937.