- Email: [email protected]
Small-scale nondestructive stress-strain and creep tests feasible during irradiation
ELSEVIER
Journal of Nuclear Materials 212-215 (1994) 1688-1692
Small-scale nondestructive stress-strain and creep tests feasible during irradiation P. Gondi, R. Montanari, A. Sili Mechanical Engineering Department, Second University of Rome, Ka della Ricerca Scientifica, 00133 Rome, Italy
Abstract By the use of suitable cylindrical indenters with flat surface and constant penetration rates, load versus penetration depth curves are obtained which show correspondence with the stress-strain curves of the tensile tests for some materials examined. After an initial linear stage, a limit load of uu is reached after which a stage of work-hardening occurs with a trend comparable to the one of the tensile test. The limit load uu z 3u,,, uY being the yielding load of the tensile test. For penetration rates = 100 km/min, the loads of the work-hadening stage above (+” tend to saturation values comparable to the loads of hardness only for some of the materials examined. With larger penetration rates, loads are reached well in excess above those of hardness. With metal-matrix composite alloys a stage with drops of load has been observed at the level of the loads of hardness. Besides those with constant penetration rates, tests with constant load have been also made on the same materials, with the results interpreted on the basis of the penetration rate effects.
1. Introduction For the optimization of the irradiation volume, miniaturization of mechanical tests on irradiated materials is of fundamental importance [1,2]. For the obtainment of indications on the fracture behaviour after irradiation, ductile to brittle transition temperatures (DBTT) and upper shelf energies (USE) are needed. Particular emphasis has been given to the small punch test (SPT) on disc-shaped samples 3 mm in diameter and with thicknesses ranging from 0.25 to 0.5 mm [3-91. In alternative, Charpy impact tests on low activation ferritic steels have been made with V-notched samples having reduced dimensions 1.5 X 1.5 X 20 mm3 [lo], disc compact tension specimens of reduced size have been used in HFIR by Elliott et al. [ll], and miniature bending fatigue samples have been studied by Rao et al. [12]. Good indications on the levels of radiation damage have been always obtained, even if shifts of the limit stresses, fracture energies and critical temperatures are 0022-3115/94/$07.00
0 1994
SSDI 0022-3115(93)E0345-A
in general experienced requiring correction factors for the comparison with data of conventional tests. However, all these tests are destructive and they are not suited to follow the evolution of radiation damage on the same sample. Moreover they are done on samples with structures which may not correspond to those of the material in use, in particular structures in zones of welding or others submitted to particular straining when in use. On account of these drawbacks and aLso in view of the possible goal of nondestructive tests made directly on the material in use, a method is considered here for which miniaturization regards not the samples but the zones of testing, with diameter u 1 mm and depths of the same order, as for hardness tests. In previous papers [13,14], indications have concentrated on the stress-strain behaviour of the material by recording load versus penetration depth of indenters with suitable geometry, different from the ones used in the common hardness tests. In the same papers, an ample bibliography can be found on previous load-penetra-
Elsevier Science B.V. All rights reserved
P. Gondi et al, /Journal of Nuclear Materials 212-215 (1994) 1688-1692 500
160 a)
1689
tion diagrams examined by various authors with the same indenters used for the hardness tests. In this paper reference is made to results obtained by the use of these new indenters with constant penetration rates and with constant load.
2. Experimental
-40
loo-
loo-
OL*L OO
070
50
100
2’5
' 200
150
+
250'
6 (Pm) 0
1
2
4
5
The present tests are based on the use of penetration by a cylindrical indenter, as sketched in Fig. 4. Results with indenters having a diameter of @ = 1 mm are recalled in this paper. For the constant penetration rate tests, the indenter advancements were recorded with precision of +l pm; the constant load penetration tests were made by progressive leaning of given weights on the indenter, with full charge occurring in times of the order of l/10 s. Constant load tests were also done with the device used for constant rate penetration, with constant load maintained through a feedback control of the penetration rate and with initial load application in the time interval of about 1 s. Tests were made on samples of Al 99.8%, Cu 99.8%, normalized AISI 1040 steel, and on the composite alloy Al-9% Si with dispersed Al,O, and SiO, particles in concentrations of the order of 10%.
3. Results
d (Pm) 0
1
5
4
2 Z(x)
2000
1500 m $ HV
b
1000
p’
4
500
@ 0
100
200
300
6 (Pm)
-i
1$010
Curves LP of load versus penetration depth of the cylindrical indenter (@ = 1 mm) for constant penetration rate of 100 um/min are shown in Figs. la-lc for the four materials, Al (Fig. la), Cu and AISI1040 steel (Fig. lb) and the composite alloy Al-9% Si with dispersed Al,O, and SiO, particles (Fig. lc). Together with the penetration curves (full lines - left hand (T ordinates and S abscissa) the stress-strain diagrams, obtained with the same materials, are drawn in comparison by the dashed lines (right hand u ordinates, E abscissa); the Vickers hardnesses HV are indicated on the left ordinate scale.
Fig. 1. Penetration curves (full lines, left hand (T ordinates, 6 abscissa) and stress-strain diagrams from tensile tests (dashed lines, right hand D ordinates, E abscissa) for Al (a), Cu and AISI 1040 steel (b), and the composite alloy Al-9% Si with dispersed Al,O, and SiO, particles. Penetration rate 100 km/min with cylindrical indenter 0 = 1 mm; strain rate in the tensile tests 10e3 i (s-l). The Vickers hardnesses (HV) of the tested materials are indicated on the left hand ordinate scale. Enlarged scale details of the penetration curves are shown in Fig. la for the initial deformation stages and in Fig. lc for a stage with drops at level of the hardness load.
1690
P. Gondi et al. /Journal
of Nuclear Materials 212-215
All the LP curves show a limit load, (do, between an initial stage of linear work-hardening and the following stage of work-hardening with the trend corresponding to u - uu a er/V with v P 3. As shown on expanded abscissa scale (Fig. la for Al as the representative example), the linear work-hardening stage starts from load oL, representing the limit for plastic deformation. Characteristics of the initial stage depend on the parallelism between the indenter basis and the sample surface. By increasing depths, the loads tend to saturation values uH which, for this penetration rate, are comparable to the hardness values for Al, Cu and Fe. With the composite alloy, loads well in excess above those of hardness are reached even with this penetration rate; in the range of the loads of hardness, a stage occurs with subsequent small drops of load, as shown in the expanded scale detail. Correspondence exists between penetration and stress-strain curves by suitable choice of the coordinates; in particular, for the deformation rates indicated, the ultimate tensile stress uR = _ au/3 and the yteld stress crY= oo/ 3; further u,_ = uY. Macroghraphic observations [13] show that in the workhardening stage above go there is evident protrusion of the material around the imprint; permanent imprints with sharp boundaries and no evident protrusion (in the first stage of workhardening) appear above uL up to uU. As mentioned, the load penetration curves of Fig. 1 have been obtained with the penetration rate of 100 pm/min. Effects of the penetration rate are illustrated in Fig. 2 for Cu, on the material taken as a representative. The penetration rate has limited effects on uo, whereas the effects on uu are relevant with three distinct stages. With the larger penetration rates, values of uu are reached well in excess above the loads of hardness. Effects of the diameter of the cylindrical indenter and of sample dimensions, as well as results of experi-
(1994) 1688-1692
400/---3-j OL------I 0 200
400
600
t
800
1000
1200
(s)
lOOO1950 8OO-F -F t
MPa
b)
1820
600.
1690 I
*O
400.
OJ
1430
0
600
1200
1600 t
2400
3000
J
3600
is)
2930 -
400. 2550 200..
2290
600
1200
1800 t
2400
3000
3600
(s)
Fig. 3. Penetration versus time diagrams in constant load penetration tests for Al (a), Cu (b) and the AISI 1040 steel (c). Cylindrical indenter @ = 1 mm, loading time l/10 s.
Fig. 2. Dependence of the limit loads mu and of the saturation loads oH on the penetration rate for Cu.
ments with subsequent stops of the penetration, are considered elsewhere [13]. Relevant recovery with load decreases after stop are observed only in the second work-hardening stage above uo. Further results, given in Figs. 3a-3c, illustrate the constant load behaviour, with the penetration versus
P. Gondi et al. /Journal of Nuclear Materials 212-215 (1994) 1688-1692 1.0
- l.OU 0
10
5
I
15
z/Q
Fig. 4. Trends of the normal (q, a,) and 45” shearing stresses (7) along the z-axis of the tested material for uniform load q applied by the cylindrical indenter sketched top right in the figure.
time curves for different
loads. For loads u < uU and in a time of some hours, depths of penetration at room temperature were within the sensitivity limit (1 km) of the displacement detector.
4. Discussion For the indenter geometry, as represented in the sketch of Fig. 4, the trends of the shearing and of the normal stresses are shown. In Fig. 4 the applied load is uniformly distributed on the contact surface. The 45” shearing stress is expressed by .j-=-
q I --
l-2u 2
3
)( 2
+(1+v)
z (pZ+*2)1’2 3
z
(p2+gy
i]
’
q = P/.rrp2 being the uniform load for applied force P, p the indenter radius and v Poisson’s ratio [15].
As visible in Fig. 4, this stress presents a maximum r* = 0.3q at depth z* z 0.6~. The limit loads uU of the penetration curves correspond thus to r* = uY, i.e. the second work-hardening stage of larger plastic deformation begins when the uY yield stress of the material is overcome at the depth z*. The first stage of linear work-hardening, from or to uo, is related on the other hand to the characteristics of rigidity of the indenter; for a rigid indenter the load q is not uniformly distributed but changes with distance r from the centre as
Conditions of cutting occur at the circular border of
1691
the indenter. The linear work-hardening stage is related to the deformation in the cylindrical core within the circular cut, until the deformation spreads over when the yielding stress is overcome at the depth of maximum shear stress. Putting L3 as the volume of spreading of the II emitted dislocation segments, the following expressions can be used: for the dislocation segment density N = n/L3, for the back-stresses Ui a n113/L and for the deformation connected with the penetration E a NSb = Sbn/L3, S being the area swept out after each emission act of the dislocations with Burgers vector b; for constant S there results ui a e1j3, in agreement with the trend of the second work-hardening stage above uo. As mentioned, the work-hardening stress increases by increasing the penetration rate; according to this model it can be explained by considering that the L spreading distance lowers with the increasing rate. Of course, this is a simplified model, that neglects the effects of the dislocation segment - pinning point distribution, dislocation sources and emission frequencies, etc. Penetration rate effects are to be taken into account in connection with the results of the constant load penetration tests. Application of full load occurs in very short times and, in consequence, reference to conditions with large penetration rates had to be made. Large constant loads can thus be applied well above the load of hardness, corresponding to the limit loads uH with low rate penetration curves experienced. At the load application, stresses higher than the backstresses cause large emission frequencies from the dislocation sources, thus the initial stage of fast deformation of these creep curves begins. When the backstresses reach levels close to the applied stress, approximately constant emission rates of the dislocation segments can be assumed and the decreases of the deformation rates appear, corresponding to the decreases of the S areas swept out after each emision act [16]. Indications on the behaviour, in the fracture of the tested materials, have been obtained only with the composite alloys when the stage of subsequent load drops refers to crack formation. The characteristics of deformation in the present tests are unfavourable to the opening and propagation of the formed cracks, so that continuous penetration occurs after the drop stage with loads well above those of hardness. It is reasonable to assume that in the composite alloy large cracks form in lengths comparable to the distances between the dispersed phase. This may explain the reasons why the stage with the drops of load has been observed only with these alloys; while searching for the corresponding stage in the other more homogenous materials, recording of the load variations with higher sensitivities and with faster responses will be used.
1692
P. Gondi et al. /Journal
of Nuclear Materials 212-215
References [l] G.E. Lucas, Metall. Trans. 21A (1990) 1105. [2] W.R. Corvin and G.E. Lucas, ASTN-STP 888 (1986) 1. [3] J.M. Baik, J. Kameda and 0. Buck, Scripta Metall. 17 (1983) 1443. [4] G. Khose, M. Ames and O.K. Harling, J. Nucl. Mater. 141-143 (1986) 513. [S] T. Misawa, T. Adachi, M. Saito and Y. Hamaguchi, J. Nucl. Mater. 150 (1987) 194. \ [6] J. MC Namey, G.E. Lucas and G.R. Odette, J. Nucl. Mater. 179-181 (1991) 429. 171A. Okada, M.L. Hamilton and F.A. Garner, J. Nucl. Mater. 179-181 (1991) 445. [8] M. Suzuki, M. Eto, K. Fukaya, Y. Nishiyama, T. Kodaira, T. Oku, M. Adachi, A. Umino, I. Takahashi, T. Misawa and Y. Hamaguchi, J. Nucl. Mater. 179-181 (1991) 441.
(1994) 1688-1692
191A. Koyama, K. Hamada and H. Matsui, J. Nucl. Mater. 179-181 (1991) 417. [lOI H. Kayano, H. Kurishita, A. Kimura, M. Narui, M. Yamazaki and Y. Suzuki, J. Nucl. Mater. 179-181 (1991) 425. Ill] C. Elliott, M. Enmark, G.E. Lucas, G.R. Odette and A. Rowcliffe, J. Nucl. Mater. 179-181 (1991) 434. 1121 G.R. Rao, A. Rowcliffe and B.A. Chin, J. Nucl. Mater. 179-181 (1991) 438. [13] P. Gondi and A. Sili, Z. Metallk. 82 (1991) 377. 1141 P. Gondi and A. Sili, Proc. XXIII Nat. Conf. on Scienza e Tecnologia Delle Superfici, A.I.M., Ancona, 1990, p. 237. [15] S. Timoshenko and J.N. Goodier, Theory of Elasticity, 3rd ed. (MC Graw-Hill, New York, 1961) p. 370. I161 P. Gondi, R. Montanari and A. Sili, Mater. Sci. Forum 94-96 (1992) 513.
Journal of Nuclear Materials 212-215 (1994) 1688-1692
Small-scale nondestructive stress-strain and creep tests feasible during irradiation P. Gondi, R. Montanari, A. Sili Mechanical Engineering Department, Second University of Rome, Ka della Ricerca Scientifica, 00133 Rome, Italy
Abstract By the use of suitable cylindrical indenters with flat surface and constant penetration rates, load versus penetration depth curves are obtained which show correspondence with the stress-strain curves of the tensile tests for some materials examined. After an initial linear stage, a limit load of uu is reached after which a stage of work-hardening occurs with a trend comparable to the one of the tensile test. The limit load uu z 3u,,, uY being the yielding load of the tensile test. For penetration rates = 100 km/min, the loads of the work-hadening stage above (+” tend to saturation values comparable to the loads of hardness only for some of the materials examined. With larger penetration rates, loads are reached well in excess above those of hardness. With metal-matrix composite alloys a stage with drops of load has been observed at the level of the loads of hardness. Besides those with constant penetration rates, tests with constant load have been also made on the same materials, with the results interpreted on the basis of the penetration rate effects.
1. Introduction For the optimization of the irradiation volume, miniaturization of mechanical tests on irradiated materials is of fundamental importance [1,2]. For the obtainment of indications on the fracture behaviour after irradiation, ductile to brittle transition temperatures (DBTT) and upper shelf energies (USE) are needed. Particular emphasis has been given to the small punch test (SPT) on disc-shaped samples 3 mm in diameter and with thicknesses ranging from 0.25 to 0.5 mm [3-91. In alternative, Charpy impact tests on low activation ferritic steels have been made with V-notched samples having reduced dimensions 1.5 X 1.5 X 20 mm3 [lo], disc compact tension specimens of reduced size have been used in HFIR by Elliott et al. [ll], and miniature bending fatigue samples have been studied by Rao et al. [12]. Good indications on the levels of radiation damage have been always obtained, even if shifts of the limit stresses, fracture energies and critical temperatures are 0022-3115/94/$07.00
0 1994
SSDI 0022-3115(93)E0345-A
in general experienced requiring correction factors for the comparison with data of conventional tests. However, all these tests are destructive and they are not suited to follow the evolution of radiation damage on the same sample. Moreover they are done on samples with structures which may not correspond to those of the material in use, in particular structures in zones of welding or others submitted to particular straining when in use. On account of these drawbacks and aLso in view of the possible goal of nondestructive tests made directly on the material in use, a method is considered here for which miniaturization regards not the samples but the zones of testing, with diameter u 1 mm and depths of the same order, as for hardness tests. In previous papers [13,14], indications have concentrated on the stress-strain behaviour of the material by recording load versus penetration depth of indenters with suitable geometry, different from the ones used in the common hardness tests. In the same papers, an ample bibliography can be found on previous load-penetra-
Elsevier Science B.V. All rights reserved
P. Gondi et al, /Journal of Nuclear Materials 212-215 (1994) 1688-1692 500
160 a)
1689
tion diagrams examined by various authors with the same indenters used for the hardness tests. In this paper reference is made to results obtained by the use of these new indenters with constant penetration rates and with constant load.
2. Experimental
-40
loo-
loo-
OL*L OO
070
50
100
2’5
' 200
150
+
250'
6 (Pm) 0
1
2
4
5
The present tests are based on the use of penetration by a cylindrical indenter, as sketched in Fig. 4. Results with indenters having a diameter of @ = 1 mm are recalled in this paper. For the constant penetration rate tests, the indenter advancements were recorded with precision of +l pm; the constant load penetration tests were made by progressive leaning of given weights on the indenter, with full charge occurring in times of the order of l/10 s. Constant load tests were also done with the device used for constant rate penetration, with constant load maintained through a feedback control of the penetration rate and with initial load application in the time interval of about 1 s. Tests were made on samples of Al 99.8%, Cu 99.8%, normalized AISI 1040 steel, and on the composite alloy Al-9% Si with dispersed Al,O, and SiO, particles in concentrations of the order of 10%.
3. Results
d (Pm) 0
1
5
4
2 Z(x)
2000
1500 m $ HV
b
1000
p’
4
500
@ 0
100
200
300
6 (Pm)
-i
1$010
Curves LP of load versus penetration depth of the cylindrical indenter (@ = 1 mm) for constant penetration rate of 100 um/min are shown in Figs. la-lc for the four materials, Al (Fig. la), Cu and AISI1040 steel (Fig. lb) and the composite alloy Al-9% Si with dispersed Al,O, and SiO, particles (Fig. lc). Together with the penetration curves (full lines - left hand (T ordinates and S abscissa) the stress-strain diagrams, obtained with the same materials, are drawn in comparison by the dashed lines (right hand u ordinates, E abscissa); the Vickers hardnesses HV are indicated on the left ordinate scale.
Fig. 1. Penetration curves (full lines, left hand (T ordinates, 6 abscissa) and stress-strain diagrams from tensile tests (dashed lines, right hand D ordinates, E abscissa) for Al (a), Cu and AISI 1040 steel (b), and the composite alloy Al-9% Si with dispersed Al,O, and SiO, particles. Penetration rate 100 km/min with cylindrical indenter 0 = 1 mm; strain rate in the tensile tests 10e3 i (s-l). The Vickers hardnesses (HV) of the tested materials are indicated on the left hand ordinate scale. Enlarged scale details of the penetration curves are shown in Fig. la for the initial deformation stages and in Fig. lc for a stage with drops at level of the hardness load.
1690
P. Gondi et al. /Journal
of Nuclear Materials 212-215
All the LP curves show a limit load, (do, between an initial stage of linear work-hardening and the following stage of work-hardening with the trend corresponding to u - uu a er/V with v P 3. As shown on expanded abscissa scale (Fig. la for Al as the representative example), the linear work-hardening stage starts from load oL, representing the limit for plastic deformation. Characteristics of the initial stage depend on the parallelism between the indenter basis and the sample surface. By increasing depths, the loads tend to saturation values uH which, for this penetration rate, are comparable to the hardness values for Al, Cu and Fe. With the composite alloy, loads well in excess above those of hardness are reached even with this penetration rate; in the range of the loads of hardness, a stage occurs with subsequent small drops of load, as shown in the expanded scale detail. Correspondence exists between penetration and stress-strain curves by suitable choice of the coordinates; in particular, for the deformation rates indicated, the ultimate tensile stress uR = _ au/3 and the yteld stress crY= oo/ 3; further u,_ = uY. Macroghraphic observations [13] show that in the workhardening stage above go there is evident protrusion of the material around the imprint; permanent imprints with sharp boundaries and no evident protrusion (in the first stage of workhardening) appear above uL up to uU. As mentioned, the load penetration curves of Fig. 1 have been obtained with the penetration rate of 100 pm/min. Effects of the penetration rate are illustrated in Fig. 2 for Cu, on the material taken as a representative. The penetration rate has limited effects on uo, whereas the effects on uu are relevant with three distinct stages. With the larger penetration rates, values of uu are reached well in excess above the loads of hardness. Effects of the diameter of the cylindrical indenter and of sample dimensions, as well as results of experi-
(1994) 1688-1692
400/---3-j OL------I 0 200
400
600
t
800
1000
1200
(s)
lOOO1950 8OO-F -F t
MPa
b)
1820
600.
1690 I
*O
400.
OJ
1430
0
600
1200
1600 t
2400
3000
J
3600
is)
2930 -
400. 2550 200..
2290
600
1200
1800 t
2400
3000
3600
(s)
Fig. 3. Penetration versus time diagrams in constant load penetration tests for Al (a), Cu (b) and the AISI 1040 steel (c). Cylindrical indenter @ = 1 mm, loading time l/10 s.
Fig. 2. Dependence of the limit loads mu and of the saturation loads oH on the penetration rate for Cu.
ments with subsequent stops of the penetration, are considered elsewhere [13]. Relevant recovery with load decreases after stop are observed only in the second work-hardening stage above uo. Further results, given in Figs. 3a-3c, illustrate the constant load behaviour, with the penetration versus
P. Gondi et al. /Journal of Nuclear Materials 212-215 (1994) 1688-1692 1.0
- l.OU 0
10
5
I
15
z/Q
Fig. 4. Trends of the normal (q, a,) and 45” shearing stresses (7) along the z-axis of the tested material for uniform load q applied by the cylindrical indenter sketched top right in the figure.
time curves for different
loads. For loads u < uU and in a time of some hours, depths of penetration at room temperature were within the sensitivity limit (1 km) of the displacement detector.
4. Discussion For the indenter geometry, as represented in the sketch of Fig. 4, the trends of the shearing and of the normal stresses are shown. In Fig. 4 the applied load is uniformly distributed on the contact surface. The 45” shearing stress is expressed by .j-=-
q I --
l-2u 2
3
)( 2
+(1+v)
z (pZ+*2)1’2 3
z
(p2+gy
i]
’
q = P/.rrp2 being the uniform load for applied force P, p the indenter radius and v Poisson’s ratio [15].
As visible in Fig. 4, this stress presents a maximum r* = 0.3q at depth z* z 0.6~. The limit loads uU of the penetration curves correspond thus to r* = uY, i.e. the second work-hardening stage of larger plastic deformation begins when the uY yield stress of the material is overcome at the depth z*. The first stage of linear work-hardening, from or to uo, is related on the other hand to the characteristics of rigidity of the indenter; for a rigid indenter the load q is not uniformly distributed but changes with distance r from the centre as
Conditions of cutting occur at the circular border of
1691
the indenter. The linear work-hardening stage is related to the deformation in the cylindrical core within the circular cut, until the deformation spreads over when the yielding stress is overcome at the depth of maximum shear stress. Putting L3 as the volume of spreading of the II emitted dislocation segments, the following expressions can be used: for the dislocation segment density N = n/L3, for the back-stresses Ui a n113/L and for the deformation connected with the penetration E a NSb = Sbn/L3, S being the area swept out after each emission act of the dislocations with Burgers vector b; for constant S there results ui a e1j3, in agreement with the trend of the second work-hardening stage above uo. As mentioned, the work-hardening stress increases by increasing the penetration rate; according to this model it can be explained by considering that the L spreading distance lowers with the increasing rate. Of course, this is a simplified model, that neglects the effects of the dislocation segment - pinning point distribution, dislocation sources and emission frequencies, etc. Penetration rate effects are to be taken into account in connection with the results of the constant load penetration tests. Application of full load occurs in very short times and, in consequence, reference to conditions with large penetration rates had to be made. Large constant loads can thus be applied well above the load of hardness, corresponding to the limit loads uH with low rate penetration curves experienced. At the load application, stresses higher than the backstresses cause large emission frequencies from the dislocation sources, thus the initial stage of fast deformation of these creep curves begins. When the backstresses reach levels close to the applied stress, approximately constant emission rates of the dislocation segments can be assumed and the decreases of the deformation rates appear, corresponding to the decreases of the S areas swept out after each emision act [16]. Indications on the behaviour, in the fracture of the tested materials, have been obtained only with the composite alloys when the stage of subsequent load drops refers to crack formation. The characteristics of deformation in the present tests are unfavourable to the opening and propagation of the formed cracks, so that continuous penetration occurs after the drop stage with loads well above those of hardness. It is reasonable to assume that in the composite alloy large cracks form in lengths comparable to the distances between the dispersed phase. This may explain the reasons why the stage with the drops of load has been observed only with these alloys; while searching for the corresponding stage in the other more homogenous materials, recording of the load variations with higher sensitivities and with faster responses will be used.
1692
P. Gondi et al. /Journal
of Nuclear Materials 212-215
References [l] G.E. Lucas, Metall. Trans. 21A (1990) 1105. [2] W.R. Corvin and G.E. Lucas, ASTN-STP 888 (1986) 1. [3] J.M. Baik, J. Kameda and 0. Buck, Scripta Metall. 17 (1983) 1443. [4] G. Khose, M. Ames and O.K. Harling, J. Nucl. Mater. 141-143 (1986) 513. [S] T. Misawa, T. Adachi, M. Saito and Y. Hamaguchi, J. Nucl. Mater. 150 (1987) 194. \ [6] J. MC Namey, G.E. Lucas and G.R. Odette, J. Nucl. Mater. 179-181 (1991) 429. 171A. Okada, M.L. Hamilton and F.A. Garner, J. Nucl. Mater. 179-181 (1991) 445. [8] M. Suzuki, M. Eto, K. Fukaya, Y. Nishiyama, T. Kodaira, T. Oku, M. Adachi, A. Umino, I. Takahashi, T. Misawa and Y. Hamaguchi, J. Nucl. Mater. 179-181 (1991) 441.
(1994) 1688-1692
191A. Koyama, K. Hamada and H. Matsui, J. Nucl. Mater. 179-181 (1991) 417. [lOI H. Kayano, H. Kurishita, A. Kimura, M. Narui, M. Yamazaki and Y. Suzuki, J. Nucl. Mater. 179-181 (1991) 425. Ill] C. Elliott, M. Enmark, G.E. Lucas, G.R. Odette and A. Rowcliffe, J. Nucl. Mater. 179-181 (1991) 434. 1121 G.R. Rao, A. Rowcliffe and B.A. Chin, J. Nucl. Mater. 179-181 (1991) 438. [13] P. Gondi and A. Sili, Z. Metallk. 82 (1991) 377. 1141 P. Gondi and A. Sili, Proc. XXIII Nat. Conf. on Scienza e Tecnologia Delle Superfici, A.I.M., Ancona, 1990, p. 237. [15] S. Timoshenko and J.N. Goodier, Theory of Elasticity, 3rd ed. (MC Graw-Hill, New York, 1961) p. 370. I161 P. Gondi, R. Montanari and A. Sili, Mater. Sci. Forum 94-96 (1992) 513.