Variation in steady state creep resistance in NiW composite eutectic alloy

Materials Science and Engineering, A I 04 (1988) 67-74

67

Variation in Steady State Creep Resistance in Ni-W Composite Eutectic Alloy S. F. DtRNFELD and Y. Z U T A

Department of Materials Engineering, Technion, Israel Institute of Technology, Haifa (Israel) (Received November 16, 1987; in revised form March 8, 1988)

Abstract

A N i - W eutectic alloy was subjected to unidirectional solidification (UDS). The post-UDS structure is an in situ composite consisting of an Ni(W) solid solution 7 matrix, reinforced with tungsten fibres or lamellae (depending on the solidification rate), and the a phase: These two phases (a + 7) are formed simultaneously from the melt in an eutectic reaction at 1495 °C. A peritectoid reaction around 970 °C induces precipitation of a third phase, fl-Ni4W, in the matrix in the form of semicoherent equiaxed particles. The dependence of the steady state creep resistance at elevated temperatures on the morphology and microstructure of the various phases was investigated. The main reinforcing effect of the a phase derives from the interaction between phase boundaries. There is a linear dependence between the steady state creep resistance and the surface area of the ct phase in both morphologies; this behaviour becomes pronounced above the peritectoid temperature, where the fl phase does not exist. A t equal surface areas, the two morphologies of the a phase provide the same contribution to the steady state creep resistance. The relative contribution of the fl phase to the steady state creep resistance decreases as the interphase spacing 2 of the a phase decreases. The creep activation energy below the peritectoid temperature is higher than the corresponding activation energy above the peritectoid temperature, and the difference between these two activation energies decreases as 2 decreases.

matrix, while the c~-W phase in the form of fibres or lamellae (Fig. 1) serves as reinforcing material [2]. ~l'he morphology of the ct phase and the interphase spacing 2 depend on the thermal gradient G at the solidification front and on the solidifica-

1. Introduction

In the course of unidirectional solidification (UDS) of the Ni-W eutectic alloy, an in situ composite is created owing to the simultaneous solidification of the two phases a + 7 [1]. In this composite the 7-Ni(W) solid solution phase is the 0921-5093/88/$3.50

Fig. 1. Transverse section of the UDS Ni-W eutectic: (a) fibrous structure, R = I cm h-~; (b) lamellar structure, R = l . 4 c m h -I.

© Elsevier Sequoia/Printed in The Netherlands

68 tion rate R. At a given constant thermal gradient, the resulting morphology is fibrous below a certain solidification rate and lamellar above it [3]. In these circumstances, although most existing theories on the mechanical behaviour of such composites at elevated temperatures are based on comparison of different eutectic systems, with different matrices and the same reinforcing phase [4], with different matrices and different reinforcing phases, or with reinforcing phases similar in nature [5, 6], the presence of both reinforcingphase morphologies in the same composite enables us to compare their influence.

the deformation modes of the specimens. Thin foils were prepared by double electropofishing in a solution of 6% H2SO 4 and 6% perchloric acid in methanol. The initial disks, 0.25 mm in thick-

/

10-4

Spec. 4" + T =900"C XL = 51am Spec. 4 • T=960"C ~.t. =Spm Spec, 4' • T=920"C XL=51Jrn Spec. 2 • T=970*C Xf=12prn Spec 3 0 T=920*CXf=31Jrn

F

~pec, 5 & T=920"C Xt.=75pm Spec. 2' V T=920=C ~f=12prn Spec. 3' "I(-T=gE~D°C X f = 3 p m

2. Experimental details

2.1. Unidirectional solidification LIDS was carried out by a Bridgman-type technique, whereby the solidification front is advanced by means of a temperature gradient shifted along the crucible which contains the liquid alloy [1]. The temperature gradient is created by means of cooling water in a doublewalled cylinder which surrounds the crucible. Because of the control of the shifting rate of the temperature gradient provided, it is possible to obtain the exact solidification rate, and through it the desired morphology and interphase spacing of the a-W phase. Examples of the resulting composite structure are presented in Fig. 1. 2.2. Creep experiments The UDS process yielded rods 12 mm diameter by 250 mm long, each of which was cut into three equal length pieces. Each of these was in turn microsliced to yield four equal disks parallel to the principal axis (growth direction), which were then shaped by machining and grinding. Measurements of the creep deformation were done by linear variable-differential transducer instrumentation. The creep experiments were performed in a Satec machine at temperatures below and above the peritectoid temperature, with an argon atmosphere to safeguard against oxidation.

Spllc. 3' x T =880*C hf = 3pro Spec. 6

D T=940*C Xf =Spm

I(T5 h=5.5

.I_ I

I00

200 500400

(oi

a" [ MPo]

Spec. 7 XT= 980"C 1

I0-3

Spec. 8 t T = I O 0 0 ° C ? X = I 2 p rn Spec. 9 A T = 1030 °C J

Spec. 10 • T= 1020"C Xf = 3pro Spec. I I o T = 1020°C ~.L=7.5 pm Spec. 12 • T = 1050"C Xf = 3pro Spec. 13 D T = I 0 2 0 * C Xt. = 5pro ",O

10-4.

2.3. Preparation of specimen Several as-grown and post-creep specimens were cut perpendicular and parallel to the growth axis, and optical metallurgy was employed for examination of the UDS structure. A JEOL 100CX scanning transmission electron microscope was used to analyse the microstructure and

10-5 (b)

i

~

i I t i i II 50 100 O tMPo]

Fig. 2. Steady state c r e e p rate vs. stress t e m p e r a t u r e a n d m o r p h o l o g y of eutectic c o m p o s i t e : (a) b e l o w the peritectoid t e m p e r a t u r e ; (b) a b o v e the peritectoid t e m p e r a t u r e .

69

ness, were polished at room temperature down to 50 # m with an electrolyte jet 1.5 mm in diameter at 120 V. The concave disks thus formed were further electropolished with the same electrolyte in a twin-jet apparatus at - 70 °C. The specimens thus prepared contained areas in which the tungsten fibres (or lamellae) and the matrix (with the/3 precipitates) were equally thinned. 3. Results and analysis

The results of the creep experiments are presented in Fig. 2. As can be seen from the linear log-log relationship between creep rate and stress, the mechanical behaviour of the composite consists of "power law creep" both below and above the peritectoid temperature. One of the most striking comparative features of these two temperature ranges is the drastic improvement in the steady state creep resistance of the composite below the peritectoid temperature at which the fl phase precipitation takes place. This abrupt change preludes direct comparison of the two creep rates of the two equally reinforced specimens under equal loads, above and below the peritectoid temperature. As loads of the order relevant to the superperitectoid ( T > 9 7 5 °C) range fall below the threshold stress for their subperitectoid ( T < 975 °C) counterparts, while those relevant to the latter range cause (when applied above the peritectoid) excessively high creep rates (which are hard to measure) or even immediate rupture. However, this feature does enable us to determine the exact peritectoid temperature under load (975 °C for a stress of 100-150 MPa). This is done by varying the test temperature up or down in steps of 1 °C and determining the transition temperature which causes an abrupt change in the creep curve. In order to estimate the relative contribution of the fl phase to the steady state creep resistance, the results were extrapolated for the fibrous and lamellar morphologies as shown in Fig. 3 (in Fig. 3(a) for specimen groups e and d, and in Fig. 3(b) for specimen groups a-c). The temperature dependence of the steady state creep resistance in the two temperature ranges (as presented in Fig. 4) enables us to extrapolate the superperitectoid results to the subperitectoid range and thus to predict the creep rate in the latter with the # phase absent (Fig. 5). Double extrapolation (in both the temperature and the stress domains) yielded Fig. 6, in which the

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I

SPECIMEN 7 dt FFIBROUSMORPHOLOGYV L kf =12pm / T =960°C iI / / i0-3 I-I

"&~oE .{0

te

SPECIMEN2

/ ~IBROUSMO~PHOLOGY~ ~iL Xf =l?_pm j

10-4

i1/

T= 970"C

1/

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10-5

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,

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SPECIMEN 13 T =1020°C SPECIMEN 4 T-~J60=C

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It

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~0-6

(b)

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Fig. 3. Comparison by extrapolation of the resistance to steady state creep below and above the peritectoid temperature: (a) specimen with identical fibrous morphology; (b) specimen with identical lamellar morphology.

70

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1250 I

1200 I

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1100

1400 I

Tk 1"550 1300 1250 f

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_ __ O- XI =12.m,0"--70 MPo 1400

1350

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1300

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Fig. 5. Prediction by extrapolation of the steady state creep rate for a eutectic without the fl phase, below the peritectoid temperature.

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IO-a 7.1

(b)

I 7.3

I 7.5

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liT E IO-4K-) ]

Fig. 4. Steady state creep rate v s . temperature: (a) below the peritectoid temperature; (b) above the peritectoid temperature.

results at the same temperature (Fig. 5) and stress (Fig. 2(a)) can be compared for specimens with and without the fl phase. The dependence of the steady state creep resistance on the a phase surface area is shown in Fig. 7 extrapolated from Fig. 2.

The phase microstructure and post-creep deformation behaviour of the same UDS specimens were investigated by scanning transmission electron microscopy in the as-grown state. A low density of dislocations was observed in the tungsten fibres after creep deformation: T= 920 °C; 2 f = 1 2 pm; g = 9 . 6 8 x 10 -5 (Fig. 8). No subboundaries were detected in either the longitudinal or the transverse cross-section even after prolonged creep. Equiaxed Ni4W semicoherent precipitates are uniformly distributed in the matrix (Fig. 9). In the matrix of post-creep specimens (at subeutectoid temperatures), relatively large plates (about 4 /~m in length) of the two phases are thus formed in the vicinity of the tungsten fibres (or lamellae); the boundary between the fl and 7 phases consists of dislocation

71 IO-Z

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10-3 O" : 150 MPo T : 920°C

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Fig. 6. Predicted steady state creep rate with and without the /3 phase, below the peritectoid temperature. 10-4

networks compensating for the incoherency between the two structures (Fig. 10). 4. Discussion

4.1. Interphase spacing2 of the a phase For the same creep stress, the steady state creep resistance of specimens with lamellar and fibrous morphology decreases as Z decreases. This fact emerges on comparison of the lamellar specimen 5 ( T = 9 2 0 ° C ; o = 1 5 1 MPa; ;tL= 7.5 #m; g=3.28X10 -5) and specimen 4' ( T = 9 2 0 ° C ; o = 1 5 1 MPa; ~L=5 #m; g= 1.03 x 10- 5) in Fig. 2(a); the creep rate of the former is three times that of the latter. For the fibrous variants, the creep rate of specimen 3 ( T = 9 2 0 ° C ; o = 1 4 0 MPa; 2f=3 ~m; g= 1.35x 10 -5) is seven times less than that of

10-5

(b)

I

OI

I

0.2

I

I

0.5 0.4 S [urn a ]

I

0.5

Fig. 7. Creep rate vs. a phase surface area: (a) below the peritectoid temperature; (b) above the peritectoid temperature.

specimen 2' ( T = 9 2 0 °C; o = 146 MPa; 2f= 12 #m; g = 9 . 6 8 x 10-5). This effect is even more pronounced above the peritectoid temperature (Fig. 2(b)). The creep rate of specimen 11 (T =1200°C; o = 8 2 MPa; 2L =7.5 #m; g= 5 X 10 --4) is about 6.2 times that of specimen

72

Fig. 8. Low density of dislocations in a fibre after creep deformation (T= 920 °C; 2f = 12 pm; g = 9.68 x 10- 5).

Fig. 10. Semicoherent phase boundary (7-fl) after creep deformation.

4.2. Influence of fl phase precipitation

Fig. 9. Equiaxed structure of fl precipitates in an as-grown specimen.

13 (T = 1200 °C; a = 8 2 MPa; 2L=5 pm; g= 8.15x 10-5). It should be noted that the lamellar spacing ratio of 1.5 for specimens 5 and 4' is the same of that of specimens 11 and 13, while the creep rate ratio of the latter specimens is about twice that of the former specimens. This higher creep rate ratio of specimens 11 and 13 arose from the absence of fl precipitates, so that the interface between the a phase and the matrix has a dominant influence. Regarding the influence of the a phase morphology, for the same stress and interphase spacing 4, the steady state creep resistance of the lamellar morphology is superior to that of its fibrous counterpart. This fact is demonstrated by the higher creep rate of the fibrous specimen 6 ( T = 940 °C; 2f = 5 pm) compared with the lamellar 4 specimen (T= 960 °C; hE = 5 t i m ) (see Fig. 2(a)), despite the fact that its creep temperature is 20 °C lower.

Above the peritectoid temperature the steady state creep resistance is significantly increased. Examples are specimen 2 ( T = 9 7 0 °C; 2~= 12 pm) and specimen 7 ( T = 7 8 0 ° C ; 2 f = 1 2 /~m), which are identically fibrous and have the same interphase spacing. According to Fig. 3(a), their predicted creep rates (under a load of 100 MPa) are 2.2x 10 -5 and 6.9× 10 -3 respectively, the creep temperatures of the two specimens being only 10 °C apart. This difference in creep rate is attributable to the presence of the fl precipitates in the matrix of specimen 2. Comparison of the lamellar specimens 4 and 13 (see Fig. 3(b)) leads to the same conclusions; in this case, however, the creep temperature difference of 60 °C at the same 2L~--"5 /~m causes a ratio of 310 between the creep rates.

4.3. Activation energies Figures 4(a) and 4(b) present the temperature dependence of the creep rate for the two morphologies in the subperitectoid and supereutectoid ranges respectively. The creep activation energies A H were computed with the aid of these graphs according to the method of Dorn and coworkers [7], using the relation

A H= R In

T , - T2

where ei and T~are the respective creep rates and temperatures at points on the graphs and R is the universal gas constant. According to this formula, the activation energies in the subperitectoid range are AH d =AH e =

5.86 x 105 J mo1-1

73

(see Fig. 4(a)) and in the supereutectoid range (see Fig. 4(b)) AHb = AHe = 5.52 X 105 J tool-l A H a = 2.93 x 105 J mol- 1 These values do not lend themselves to simple physical interpretation, nor can they be connected directly to the diffusion activation energy, but some comparisons are possible. At elevated temperatures (with fl precipitates absent) the value for specimen a (;if= 12/~m) is much lower than that for specimen b (;if= 3 /~m) and specimen c (;iL= 7.5 /~m). In contrast, A H b and A H c are almost equal. These data indicate higher temperature sensitivity for the specimen with the coarsest morphology, i.e. specimen a (;if=12 /~m), where there is a contribution of the y matrix to the creep process. For the finer morphologies, AHb and A H c (below the peritectoid temperature) are seen to be a little higher than AHa and A He, owing to the presence of the fl precipitates. 4.4. Contribution of different morphologies to the steady state creep resistance In order to assess the relative contributions of the different morphologies, the subperitectoid creep rate values with fl precipitates present should be compared with the predicted values in this temperature range for a composite without fl precipitation. In principle, the comparison could be realized by extrapolation of lines a, b and c (Fig. 4(b)) to the subperitectoid range (see Fig. 5), but in practice this is impossible as the stresses involved are too low to induce creep. Accordingly, parallel extrapolation of the g values in Fig. 2(b) must be carried out for the range of low stresses (see Fig. 6). The results of the comparison, and the predicted g values, are presented in Table 1. The table shows that the relative contribution of the fl phase precipitation to the steady state creep resistance is much larger in the coarse fibrous case (;if= 12/~m) than in the fine TABLE 1

Morphology

Fibrous Fibrous Fibrous Fibrous Larnellar

fibrous case (;if= 3 pm) and the lamellar case (;iL = 7.5 #m). Apparently, the composite is characterized by an inverse correlation between the interfibre and interlamellar spacing and the steady state creep resistance. Regarding the correlation between the surface areas Sf and SL of the a phase and the spacings ;if and ;ie, analysis yielded the following relation for a volume element [8]: 2

SL'~-~L

(2) 0.95 Sf-

2f

The relationship between creep rate and surface area for the two temperature ranges is presented in Fig. 7. This log-linear relationship can be formulated as In g = In"A1- mS

(3)

where g is the steady state creep rate at a definite temperature and load, A~ is a constant dependent on temperature and load, m is an independent constant and S is the surface area of fibres or lamellae. The divergent value for ;if= 12 ,urn in Fig. 7(a) can be explained by the fact that, with such a coarse morphology, the creep rate is more strongly dependent on the fl phase precipitation and reduces the relative contribution of the a phase surface area to the steady state creep resistance. Similar relations between the creep behaviour and the surface area in fibrous composites were observed by Kossowsky [9] and Bullock et al. [10]. The above indicates that the creep behaviour is independent of the morphology so long as the surface area is the same, contrary to the findings of Brenian et al. [11] and Bibring and Rabinovitch [12], who claimed that the fibrous morphology is superior.

Relation of predicted creep ratio (below predicted temperature) )~f or 2 L

Predicted g,

Predicted g,

(/~m)

a+7+fl (b-')

a+y (h-')

Predicted ratio of g

Ag notation in Fig. 6

12 12 3 3 7.5

4.6 x 10 6 1.0xl0 6 1 . 0 x 10 7 5.8 x 10 -~ 8 . 2 x 10 -8

1 . 4 x 10 -3 3.0x10 4 2.5 x 10 -6 1.3 x 10 -7 2.3 x 10 6

304 300 25 22 28

AgA AgB A ~:E AgD Ag v

74

5. Conclusions

(1) A n Ni-W eutectic alloy was subjected to UDS. A peritectoid reaction at around 970 °C results in formation of semicoherent Ni4W precipitates, the fl phase. (2) The dependence of the steady state creep resistance at elevated temperatures on the morphology and microstructure of the various phases was investigated. The following features were found. (a) The relative contribution of the fl phase to the steady state creep resistance decreases as the interphase spacing of the a phase decreases. (b) The dependence of the steady state creep resistance on the a interphase spacing becomes more pronounced above the peritectoid temperature, where the fl phase does not exist. (c) The behaviour of the Ni-W eutectic composite is consistent with power law creep. (d) The creep activation energy for fine morphologies is almost the same below and above the peritectoid temperature: AH~5.86 x 105 J mol- 1. (e) The main reinforcing effect of the a phase derives from the phase interaction. There is linear dependence between the steady state creep resistance and the surface area of the a phase in both morphologies.

(f) The two morphologies contribute differently to the steady state creep resistance owing to the different surface areas of the two phases (at equal interphase spacings). References 1 M. Bamberger, S. E Dirnfeld and Y. Zuta, J. Cryst. Growth, 73(1985) 142. 2 S. E Dirnfeld and D. Shechtman, Metall. Trans. A, 16 (1985) 1187. 3 S. F. Dirnfeld, Y. Zuta and D. Schwam, Proc. Int. Conf. on Speciality Steels and Hard Materials, Pretoria, Pergamon, Oxford, 1982, p. 345. 4 L.D. Livingstone, J. Appl. Phys., 40(1970) 192. 5 K. N. Street, F. C. St. John and G. Piatti, J. Inst. Met., 95 (1967) 326. 6 W. Kurz and B. Lux, Metall. Trans., 2(1971) 329. 7 J. E. Bird, A. K. Mukherjee and J. E. Dorn, Proc. Int. Conf. on Quantitative Relations Between Properties and Microstructure, Haifa, 1969, Universities Press, Jerusalem, 1969, p. 255. 8 S. F. Dirnfeld and Y. Zuta, in H. Lilholt and R. Talreja (eds.), Fatigue and Creep of Composite Materials, Proc. 3rd Riso Int. Symp. on Metallurgy and Materials Science, 1982, Riso National Laboratory, Ris~, 1982, p. 205. 9 R. Kossowsky, Metall. Trans., 1 (1970) 1905. 10 E. Bullock, M. McLean and D. E. Miles, Acta Metall., 25 (1977) 333. 11 E. M. Brenian, E. R. Thompson and W. K. Tice, Metall. Trans., 3(1972) 132. 12 H. Bibring and M. Rabinovitch, Int. Symp. on Composites Grown in Situ, Eindhoven, 1971.