ML20147B566
| ML20147B566 | |
| Person / Time | |
|---|---|
| Site: | Point Beach, 05000000 |
| Issue date: | 06/15/1973 |
| From: | Perrin J, Scotti V, Sheckerd D Battelle Memorial Institute, COLUMBUS LABORATORIES |
| To: | |
| Shared Package | |
| ML20147B514 | List: |
| References | |
| FOIA-88-45 NUDOCS 8803020122 | |
| Download: ML20147B566 (115) | |
Text
FINAL REPORT on POINT BEACH NUCLEAR PLANT UNIT NO. 1 PRESSURE VESSEL SURVEILLANCE PROGRAM:
EVALUATION OF CAPSULE V to WISCONSIN ELECIRIC POWER COMPAb7 June 15, 1973 I
by 4
J. S. Perrin, J. W. Sheckherd, D. R. Farmelo, and L. M. Lowry Approved by BATTELLE Columbus Laboratories 505 King Avenue Colu:nbus, Ohio 43201 8803020122 880226 PDR FOIA gl CONNOR 88-45 PDR
4 TABLE OF CONTENTS E.*Lt
SUMMARY
1 INTRODUCTION 2
BACKCROUND 5
CAPSULE RECOVERY AND DISASSDiBLY 10 SAMPLE PREPARATION 13 Pressure Vessel Material 13 Correlation Monitor Material 13 EXPERIMENTAL PROCEDURES 14 Dosimeter and Thermal Monitor Examination 14 Impact Tests 16 Tensile Tests 19 Fracture Toughness Tests 22 RESULTS AND DISCUSSION 26 Dosimeter and Thermal Monitor F.xamination 26 Impact Tests 37 Tensile Tests 56 Fracture Toughness. Tests 72 CONCLUSIONS 80 REFERENCES 31 4
APPENDIX A 1
PRESSURE VESSEL MATERIAL A-1 APPENDIX B CORRELATION MONITOR MATERIAL B-1
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APPENDIX C f
INSTRUMENTED CHARPY EXAMINATION C-1 l
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l POINT BEACH NUCLEAR PLAhT UNIT NO.1 PRESSURE VESSEL SURVEILLANCE PROGRAM: EVALUATION OF CAPSULE V by J. S. Perrin, J. W. Sheckherd, D. R. Farmelo, and L. M. Lowry EHMARY The irradiation conditions and the irradiation-induced changes in mechanical properties of the Point Beach Nuclear Plant Unit No.1 reactor pressure vessel (SA302 Grade B) have been determined from evaluation of specimens contained in surveillance Capsule V.
This capsule contained base
, metal, heat-affected zone metal, and weld metal specimens.
The capsule was removed after 1,49 equivalent full-power years of operation.
The irradiat!cn te=perature did not exceed 590 F, and the e,apsule obtained a fluence of
_3.58 x 10 nyt (>l Mev).
The measured changes in nil ductility transition temperature (NDTT) for the three materials were consistent with those observed for other surveillance programs involving similar materials and irradiation conditions.
A Telatively well defined trend band for change in NDTT with increasing exposure to fast neutrons was determined from the results of this program and those of the other programs. The upper-bound curve of this trend band indicates the projected 32 equivalent full power year (end of life) change in IUrrr is in good agreement with the original predicted shif t.
The yield and ultimate tensile strengths of the materials examined increased as expected.
The present investigation shows that the pressure vessel base metal, the veld metal, and the heat-affected zone metal mechanical properties are changing with Avradiation in a manner in agreement with the changes expected when the pressure vessel was constructed.
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i INTRODUCTION This report presents' the results of the examination of Capsula V, the first capsule of the continuing surveillance program for monitoring i
i the effects of neutron irradiation on the SA302 Grade B Point Beach Unit No.1 reactor pressure-vessel material under actual operating conditions.
This report contains experimental procedures, results, and discussion relating to the investigation.
j Radiation damage studies initiated during the early days of nuclear power-reactor development revealed the deleterious effects of high
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energy neutrons upon the notch ductility of reactor vessel steels. The effect was characterized by a rapid rise in the nil ductility transition temperature (NDIT) with increasing neutron exposure. In addition, the tensile properties show a significant loss of uniform elongation and reduction of area with increasing neutron exposure.
Sufficient data on the effects of radiation on the mechanical J
properties of reactor pressure-vessel steel are now available to indicate i
j the type and relative magnitude of property changes to be encountered during
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the expected lifetime of the reactor structure. This information is an j
j integral part of the design basis for a nuclear reactor.
At the time of
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startup of the Point Beach Unit No.1 reactor, the pressure-vessel asterials f
were of sufficient quality to ensure that the expected.adiation-induced
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changes in mechanical properties would permit continued safe operation of i
the reactor even at the projected end-of-life period of 40 years of operation at an 80 percent load factor (32 equivalent full power years).
During this period the reactor operating limitation curves (i.e., pressure and temperature) vill be periodically adjusted to incorporate the projected changes in mechanical properties.
To further ensure the continued safe operations of the plant, a
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reactor-vessel radiation-surveillance program is being conducted. The I
primary purpose of this program is to evaluate the specific changes in the f
t mechanical properties of the pressure-vessel materials under the actual
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service conditions (neutron fluence, time, and temperature) of the reactor
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plant. It is known that the magnitude and relationships of the property changes are functions of the specific material composition and metallurgical 1
I condition; the amount, rate, and energy spectrum of the radiation; i
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and the exposure temperature (1-7)*
1he surveillance program in designed to I
provide information for determining whether the reactor pressure-vessel l-operating limitations are indeed conservative, as is expected.
l The surveillance program for the Point Beach Nuclear Plant Unit j
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No. I was designed and reconumended by the Westinghouse Electric Corporation
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t and is based on ASIM E 185, "Surveillance Tests on Structural Materials in I
Nuclear Reactors"I }
The details of this program and the preirradiation mechanical properties of the materials are presented in Reference (9). Prior to startup, six capsules containing tensile, Charpy V-notch, and WOL fracture-mechanics specimens of the pressure-vessel materials were installed in the f
reactor. The capsules were located between the thermal shield and the vessel l
6 vall. In addition to these mechanical-property test specimens, the capsules t
contain thermal-monitor and neutron-fluence specimens for evaluation of the specific temperature and radiation exposure conditions of the specimens.
l The particular exposure condition variables evaluated are the total
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integrated fast fluence of the capsule and the maximum temperature encountered l
4 by the specimens during the exposure period.
The temperature history of the i
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surveillance capsule is fairly representative of that encountered by the 1
j pressure-vessel vall. However, the capsule is a finite distance from the reactor pressure-vessel wall and, therefore, the capsula received an i
accelerated fluence as compared to the vessel vall.
j The most essential mechanical properties evaluated by the test specimens in the surveillance capsule are the ductile-to-brittle fracture f
I transition temperature and the conventional tensile-strength and ductility l
l values. In this context, essential refers to those requirements of the current
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methods for establishing pressure-temperature operating limitations of the l
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reactor pressure vessel and not necessarily the more advanced and/or f
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sophisticated methods for setting safe operating limitations of a structure.
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- References at end of text.
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4 An essential requirement of the mechanical property measurements is that J
they be made on representative material. For this surveillance program, the capsules contain test specimens of the SA302 Grade 3 re2ctor-vessel l
steel from two 6-3/4 in, thick shell plates from the vessel intermediate
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l and lower shell courses adjacent to the core region, and also veld metal and f
i heat-affected zone (MA2) metal. The thermal history of the material used to l
fabricate test specimens is as identical as possible to that received by the reactor pressure vessel during fabrication, with the exception that the t
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specimen post-weld heat treatment has been simulated.
In addition to the l
reactor materials, test specimens of a epecially prepared correlation material j
(SA302 Grade B) made available by Subcotanittee II of ASTM Committee E10 were also contained in the capsule. The data obtained from evaluations of the correlation caterial provide a valuable link with the surveillance programs of other nuclear-reactor pressure vessels.
a An advanced materials-evaluation technique was used during the course of this exa:nication. This advanced technique was instrumentation of the Charpy test to obtain additional data during Charpy testing. The additional I
information obtained during this testing is presented in Appendix C, "Instrumented Charpy Examination".
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5 BACKCRO WD The purpose of this section is to supply background information on the general features of radiation damage and the current requirements for assessment of the fracture-safe operating criteria for reactor pressure-vessel steel.
The overall effects of fast neutron irradiation on the mechanical properties of low-alloy ferritic pressure-vessel steel are well documented in II'0) the literature The minimum integrated fast neutron exposure necessary to produce changes in the mechanical properties of this class of steel is somewhat dependent on the reactor operating temperature.
In general, a fluence greater than 10" nyt (>l Nev) is required for temperatures between 400 and 500 F, and for temperatures greater than 500 F a fluence greater 8
than 10 nyt (>l Mev) is required to produce a measurable change in mechanical prope rtie s. The operating temperature of the Point Beach Unit No. I reactor pressure vessel is approximately 550 F, and t.his teeperature is in the range where large exposure times are required to produce measurable mechanical-property changes. The amount of radiation damage is also strongly dependent on minor changes in the chemical composition of these ferritic steels (10,W This is very important when co=parisons are made between the properties of base metal and weld metal.
The chemical composition of the base metal is well controlled and characterized. However, the current indestry practice is such that the cceposition of the heat-affected-zone and weld metal is not as well characterized as the base metal.
Since irradiation increases the yield and tensile strength values (at the expense of a reduction in ductility), the major emphasis in a surveillance program is devoted to considerations of tha ductile-to-brittle transition te=perature.
This property is usually determined by tests with notched specimens impact loaded to fracture using the Charpy V-notch impact test. The general effect of fast neutron irradiation is to produce an increase in this transition temperature. There is considerable documentation of the effects of neutron irradiation on the notch ductility properties of cocoon structural materials (
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6 Investigations at the Naval Research Laboratory are responsible for the development of the widely used nil ductility transition temperature concept for establishing fracture-safe operating limits of ferritic steel structures. The NDTT is a function of the fundamental defomation and fracture properties of the material and is defined by the Drop-Weight Test, as discussed in ASTM E208, "Conducting Drop-Weight Test to Determine Nil.
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Ductility Transition Temperature of Ferritic Steels" In this test, the NDTT is defined as the temperature at which the drop-weight specimen is broken in a series of tests in which duplicate no-break refers to the type of specimen fracture where the crack is arrested before traversing the width of the specimen. Therefore, at temperatures above the NDTT the specimens are ductile and at temperatures below the NDTT the specimens are brittle.
This NDTT concept was developed by Pellini(16) and has proven a good quali-tative tool for avoiding brittle fracture of ferritic steel structures.
The NDTT is used as a primary index temperature for the Fracture Analysis D' ngram (FAD) interpretive approach to brittle fracture prevention on i
which the "NDTT +60F" criterion is basedII6).
The standard drop-veight test specimen is prohibitively large for inclusion in most radiation-surveillance programs. The Charpy V-notch specimen is relatively small and widely used; therefore, it has been adopted as a primary specimen for radiation surveillance studies. T9 be consistent with l
l the concepts required for the FAD, a Charpy correlation energy or "fix" is l
f normally detemined for a particular steel to index the NDTT. For the SA302 Grade B steel the "fix" temperature which corresponds to the NDTT is, that at which a correlation energy of 30 f t-lb occurs for the standard Charpy When referring to an NDTT measured by Charpy V-notch V-notch imps.ct test test techniques, it is understood that this implies a prior knowledge of the correlation energy.
Fracture orientation is another concept that should be considered when assessing the fracture-safe operating criteria for a reactor pressure vessel from the results of surveillance-specimen evaluations. There are weak i
and strong fracture directions in most ferritic materials.
In this context, direction refers to the plane and path of the crack. The transition energy curves fo,r these two directions will have different upper-shelf energy values,
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j with that for the weak direction being the lower. At the 30 f t-1b fix temperature it is difficult to discriminate between different fracture directions. Therefore, the interest in difference in fracture directions is directed toward considerations of the upper shelf energy values.
The reactor environment may alter the pressure-vessel material
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properties to a condition where NDTT concepts of fracture-safe operation no longer apply, by causing the upper shelf energy to drop to a sufficiently
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low value. This type of behavior is known as low-energy tear fracture, where only a relatively small amount of energy (i.e., usually comparable to j
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that associated with brittle low. temperature fractures) is required for extensive propagation of a ductile fracture.
This behavior is typical of
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4 some high-strength steels, and the radiation damage caused by the reactor 4
environment can be regarded as promoting a strength transition of fracture properties.
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The point at which NDTI concepts no longer apply is not well defined. However, some investigators (17) identify this point as that where f
1 the upper-shelf energy is equal to or less than 50 ft Ib for a transition I
temperature curve determined by Charpy impact tests. In addition to a shif ting of the transition energy curve upward in temperature, radiation damage i
decreases the difference between the upper and lower shelves by essentially
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lowering the upper-shelf energy, i
The reactor pressure vessel under consideration was designed and j
fabricated so that the operating stresses imposed upon the weak direction of t
the material are lower than those for the strong direction. In addition, the design is such that vessel penetrations having complex stress states are sot t
subjected to radiation levels which would appreciably alter the material properties, i
In addition to the transition teroperature approach, there is the f
fracture sechanics approach to studying the problem of brittle nacerial f
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failure Ite fracture-mechan ws approach relates the stress field at I
l the tip of a crack to the requirements for brittle fracture. Here, the conditions for brittle fracture are defined from knowledge of the stress at j,
the crack tip caused by stress on the bulk structure, the resistance of the j
sacerial to crack propagation, and the sise and severity of a defect capable ef initia, ting fracture.
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The stress field in the materisi near the crack tip is defined by a single parameter known as the stress intensity factor K.
This parameter is a function of the geometry of the bulk structure, the geometry and location of the crack, and the distribution and magnitude of external I
loads on the bulk structure. Therefore, if a component is desigt.ed so that the geometries of the structure and defect are known, the stress intensity at the crack tip can be obtained from the applied stress and the size of the defect. The value of K has been determined for many practical laboratory cc,nfigurations.
The wedge open losding or WOL specimen design used in this work has been evaluated to the point 'iner the relation between applied load and stress intensity is well defined.
l The brittle fracture of a material occurs when the magnitude of the stresses at the crack tip exceeds some critical value. When the material is loaded in tension with the flav perpendicular to the direction of the load I
and there is limited local plasticity, then this critical value of X is the 4
j plane-strain fracture toughness, K The value of K at a particular load g.
g rate and temperature can be considered an intrinsic material property which is a measure of the resistance of a material to brittle failure in the presence l,
of a stressed crack.
The value of K, for a material is affected by metallurgical and g
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.sechanical variables.
In addition to the effects of load rata and temperature,
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previously mentioned, the measurement of K is hindered by the effects of R
specimen size. The condition of plane strain is essential, for the assumptions l
of elastic behavior are inhecent in the fracture-mechanics approach.
If there is a great deal of local plastic deformation around the crack, the measured critical value of X is not one of plane strain and therefore it is not a valid i
be measurement. There is a maximum plastic. zone size that can be tolerated 1
without affecting the validity of the measured value of K For the specimen ye.
geometry used in this work, the plastic.none size can be calculated from the r
measured value of K, and yield strength. The thickness (B) of the WOL 7
specimen used must be 22.5 (Kg/c,)2, where e, is the yield stress, s is y
y termed the ASm validity criterion for WOL fracture-toughness specimens 3
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9 From a consideration of this cri:acion it can be seen that once a specimen 3
size is determined (fixing B) the test temperature alone will determine i
l whether a valid K, can be measured.
Thus, the results must be anticipated g
before the test temperature is selected, and the yield strength must be known as a function of temperature.
It can be seen from the validity requirements that evaluation of materials having high K, values and low yield strengths requires very thick 7
l specimens. When the use of large specimens is not practical, the test temperature must be decreased so that an adequate reduction in K, and 7
attendant increase in yield strength will permit satisf action of the validity
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requirements, i
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10 CAPSULE RECOVERY AND DISASSIMBLY Batte11e's Columbus Laboratories (BCL) personnel vent to the Point Beach Nuclear Plant Unit No I to pick up the surveillance capsule assembly. They brought a pool-side jib crane, a specialized underwater cutting tool, and a shipping cask. The cutting head of the underwater cutting tool is a mild steel casting.
The head had been sand blasted, copper plated, and then nickel plated to prevent it from rusting and thereby contaminating the pool water.
To further avoid contamination, pool water was used in the line leading from the pump inter.sifier unit to the cutting head. The specialized underwater cutting tool is shova in Figure 1.
The tool is shown positioned on top of the shipping cask in Figure 2.
The capsule assembly had an overall length of 131 inches.
Point Beach personnel removed the capsule assembly from the pressure vessel and transferred it underwater in a canal to the spent fuel pool. The upper lid and lover drain lid were removed from the shipping cask.
Using an overhead crane, the cask was then raised from the receiving area, moved to a positinn over the spent fuel pool, and lowered into the pool so that the bottom end of the cask was resting on the floor of the pool. The bridge crane was then P
used to position the capsule and attached lead tube such that the capsule was in the cask except for the lead tube and about six inches of the capsule.
The cutting tool was lowered into the pool usf og a stainless steel c5ble attached to the pool-side jib crane.
The cutter was guided into position using the stainit.ss steel pipe line leading to the cutting head.
Binoculars were used to determine the position of the cutting head. The lead tube was separated from the capsule assembly by making a cut about 2 inches above the top end of the capsule.
The lead tube was placed into the cask along side the capsule.
The cask was raised from the pool and its exterior was thoroughly l
rinsed with water. The water inside the cask was allowed to drain into the pool. The overhead crane was then used to lower the cask to the decontamination area. The cask was decontaminated to the level of removable i
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FIGURE 1.
SPECIALIZED UNDERWATER CUTTING TOOL.
This tool is used to cut the capsule lead tube to separate it from the capsule.
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CUTTING TOOL PLACED ON TOP OF SHIPPING CASK 1
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j contanination required for shipping, 2200 disintegrations /100 cm / min By 2
and 220 disintegrations /100 cm / min a.
The cask was then shipped to the j
q DCL Hot Laboratory Tacility by commercial carrier.
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Upon arrival at BCL, the cask was placed in a hot cell. The l
capsule and lead tube sections were then removed from the cask. The capsule was examined to confirm the cut separating the lead tube from the i
capsule was in the lead tube, thereby maintaining the water-tight integrity of the capsule. Visual examination showed the capsule to be a dark gray color. The capsule had a slight bow, but subsequent examination show(d that the bowing had not damaged the specimens in the capsule.
The specimens were removed from the capsule and inventoried. The appearance of the specimens was a dark gray color. Before testing, the mechanical property specimens were desned in the following manner: washed in I
Chlorothene, washed in Radiacwash and distilled water, rinsed twice in
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distilled water, and rinsed in reagent grade alcohol.
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SAMPLE PREPARATION 1
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pressure vesset Material Babcock and Wilcox supplied the SA302 Grade 3 reactor pressure 4
vessel material frca plates C1423 and A9811. These two plates were used in 1
j ghe lower and intermediate shell courses of the vessel.
Babcock and Wilcox
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also supplied a weldment which joined the two shell plates. Appendix A contains the chemical snalyses and thermal treatment history of the plate l
materials.
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i esterial from each shell plate was beat-treated with the shells.
.n' All test
.sens were machined from the 1/4 thickness location of the plate after peru raing a simulated stress-relieving treatmed.
The test specimens
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represent material taken at least one plate thickness (6 3/'. inches) from the l
t quenched edges of the plate.
Specie ns were machined fres veld and heat-affected none metal from a stresr tet tevid weldment joining the twc she11 i
plate materials. All heat-af fected av..a specimens were obtained from the I
weld-heat affected zone of plate A9811.
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. The axis of the notch of the Charpy V notch impact specimens was a
machined perpendicular to the major surfaces of the plate. The longitudinal l
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axis of, the impact specimens was parallel to the rolling direction of the plate.
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All teattle specimens were machined with the longitudinal axis of the specimen j
parallel to the rolling direction of the plate.
All Wot test specimens were j
machined with the simulated crack in the specimen perpendicular to the 1
rolling direction and the major surfaces of the plate.
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Correlation Monitor Material t
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The SA302 Grade B material for the correlation monitors was supplied f
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by the U.S. Steel Corporation from 6-inch-thick plate. Appendix 5 contains i
the chemical analyses and thermal treatment history of the correlation monitor materials.
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14 EXPERIMENTAL PROCEDURES This section describes the procedures employed in the testing of the impact, tensile, and k'OL specimens. Also included are the procedures used to examine the dosimeters and thermal monitors. All testing and evaluations were performed at Batte11e's Columbus Laboratories.
Dosimeter and Thermal Monitor Examination The capsule contained two kinds of low-melting-point eutectic alloy thermal monitor wires for determination of the maximum temperature attained by the test specimens during irradiation. These thermal-monitor wires were sealed in Pyrex tubes and inserted in spacers in the capsule.
During capsule disassembly the thermal monitor wires were removed from the spacers and Pyrex tubes for visual examination.
The capsule contained dosimeters of copper, nickel, cadmium-shielded aluminum-cobalt alloy, unshielded aluminum-cobalt alloy, neptunium 237 and uranium 238. In addition, the tect.anical test specimens provided material for iron.dostmeters. The reactions used for the dosimetry calculations were as follows:
'Fe (n,p)
Mn Iron:
Nickel:
Ni (n.p) 58 58 co 0
O Copper:
Cu (n,c)
Co 9Co (n,y) 60 Cobalt:
Co 238U (n,f) 137Cs Uranium :
Np (n,f)
Cs All dostmeter samples were analyzed except two.
One bare cobalt wire was not recovered during capsule disassembly, and one cadmium-shielded cobalt wire was not analyzed due to difficulty in removing the wire from the cadmium shield.
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15 After removal from the capsule, the individual samples were placed in vials for transfer to the radiochemistry laboratory.
Radiation readings at 1 meter and on contact were recorded. The nickel, copper, and cobalt wires were decontaminated by wiping with dilute acid, 238 distilled water, and reagent grade acetone. The iron samples, and U
and Np capsules were wiped with dilute acid and distilled water to remove niajor contamination and then cleaned ultrasonically in a solution of Radiac and water.
The pure copper wires were weighed to i0.0001 g, and the activation product ( Co) intensity was determined directly by ga=ma ray spectrometry without dissolving the samples.
The iron, nickel, and cobalt (Al-0.15 percent Co) samples were weighed, dissolved in a solution of hcl and HNO,
3 diluted to appropriate volumes, and 1 mi samples were taken for gamma counting.
8 237 0 and Np capsules were opened in an alpha radiation containment box by specially prepared tools used to grip the small 1/4 in, diameter x 3/8 in.
, long cylinders and cut of f the tops.
The tool used for cutting off the tops O
37 was a modified tubing cutter. The U and Np were present in the form of oxide powders. The two samples were poured into small cared primary contain-ment vials and then into clean tared secondary vials for weighing to
- 0.0001 g on an analytical balance.
They were dissolved in 83 HNO3 (U 0 ) and 38 8M H SO -0.13 NaBr03 (Np0 ), and diluted to appropriate volumes. 137Cs analyses 2 4 2
were performed in duplicate af ter purification by the chloroplatinate method.
For assurance of complete fission product decontamination, Zr and Ru holdback carriers were employed, and an extra scavenge precipitation step was performed.
All the activation products were analyzed by ganusa-ray spectrometry utilizing a 3 in, diameter x 3 in. long NaI (T1) scintillation crystal detector and model ST 400 K Tullamore transistorized 400 channel analyzer (Victoreen Instrument Co.) capable of 8.5 percent resolutivn FWHM (full width half maximum) at the 0.663 Mev Cs 137"Ba gamma ray energy level.
1 The 60 137 54 Co and 0s samples were counted directly against NBS standards.
The Mn and 58Co activities were obtained from comparison with theoretical efficiency curves prepared from NBS standards.
The procedurec used in the evaluation of the dosimetry samples followed the appropriate ASIM recommendations ( 0-26) i 4
l 16 Impact Tests The impact tests were performed on a standard Wiedemann-Baldwin testing machine in accordance with the recommendations of the pertinent ASTM standard (
)
The accuracy of the machine was verified on November 23, 1972 with standards purchased from the U.S. Army Materials Research Agency. The results are given in Table 1.
TABLE 1.
CALIBRATION DATA FOR THE BCL HOT LABORATORY CHARPY IMPACT MACHINE TESTED ON NOVDiBER 23, 1972 Average BCL Energy, Standard Energy (8)
Variation Group f t-lb ft-lb Actual Allowed Low Energy 13.0 12.8
+0.2 f t-lb
- 1.0 f t-lb Medium Energy 41.7 42.6
-2.1 percentt5.0 percent High Energy 68.1 69.4
-1.9 percent *5.0 percent (a) Established by U.S. Army Materials and Mechanics Research Center The design for the specimens used for all impset tests is shown in Figure 3.
ASTM test procedures for specimen temperature control were utilized.
The low temperature bath consisted of agitated methyl alcohol cooled with additions of liquid nitrogen.
The container was a Dewar flask which contained a grid to keep the specimens at least 1 in from the bottom.
The height of the bath was enough to keep a minimum of 1 in of liquid over the specimens.
The Charpy specimens were held at temperature for a minimum of at least the ASIM recommended time.
The tests above room temperature were conducted in a similar manner except that a metal container with a liquid bath was used.
The bath used for temperatures from 70 to 212 F was water, and the bath used for temperatures above 212 F was oil. The baths were heated to temperature using a hot plate.
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18 The specimens were manually transferred from the temperature bath to the anvil of the impact machine by means of tongs that had also been brought to temperature in the bath.
The specimens were removed from the bath and impacted in less than 5 sec. The energy required to break the specimens was recorded and plotted as a function of test temperature as the testing proceeded. Selection of the test temperatures was based primarily on finding the temperature corresponding to a specimen absorbed energy of 30 f t-lb.
The correlation between Charpy V-notch test results and the nil ductility transition temperature (NDTT)* for SA302 Grade B steel has been shown to be 30 f t-lb(9)
Lateral expansion was determined from measurements made with a vernier caliper.
Fracture appearance was estimated from observation of the fracture surface, and comparing the appearance of the specimen to an ASTM fracture appearance chart ( 0)
Tne capsule contained a total of 48 Charpy V-notch specimens.
Base metal spe~cimens were from plates C1423 and A9811, which were used in the lower and intermediate shell courses of the vessel.
Heat-affected zone and weld metal specimens were from a weldment which joined the two shell plates.
Correlation monitor Charpy specimens were from material furnished by the U.S. Steel Corporation through Subcommittee II of ASTM Committee E10 on Radioisotopes and Radiation Effects, i
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_ Tensile Tests The design of the tensile specimens is shown in Figure 4; the gage section has a nominal 0.250-in. diameter and a nominal 1.000-in. length.
The tensile tests were conducted on a screw-driven Instron testing machine having a 20,000-lb capacity. A crosshead speed of 0.02 in, per min was used.
The defomation of the specimen was measured using a strain gage extensometer.
The strain gage unit senses the differential movement of two extensometer extension ams attached to the specimen gage length 1 in, apart. The extension arms are required for thermal protection of the strain gage unit during the elevated temperature tests.
Figure 5 shows the extensometer extension ams and strain p, age assembly used for tensile testing. A tensile specimen is shown at the top of the figure next to the region of the extension arms where the specimen is loaded for testing. The strain gage unit is shown at the bottom of the figure next to the region of the extensometer arms where the unit is attached during testing.
The extensometer was calibrated before testing using an Instron high-magnification drum-type extensometer calibrator.
Elevated temperature tensile tests were conducted using a three-zone split furnace.
An unirradiated tensile specimen with several thermocouples directly attached to the gage section was used to determine the optimum pcwer input to each furnace zore before testing irradiated specimens. The irradiated tensile specimens were tested at 550 F.
The specimens were held at temperature before testing to stabilize the temperature. Temperature was monitored using a Chromel-Alumel themocouple in direct contact with the gage section of the specimen. Temperature was controlled within 3 F.
Lov temperature tests were made using a liquid nitrogen spray cryostat. The cryostat was of a split clamshell type. Liquid-nitrogen vapor was used to cool the specimen and the containment chambers. The liquid nitrogen flow rate was adjusted by means of a valve in the gas line, which allowed the specimen temperature to be closely controlled during testing. Tests were run at -180 F and -110 F, with temperature controlled within *2 F during testing.
Temperature was measured using a copper-constantan thermocouple in contact with the specimen gage section.
The load-extension data were recorded on the testing machini strip chart.
The yield strength, ultimate tensile strength, and total elongation were determined from these charts. The reduction in area was determined from specimen measurements made using a vernier caliper.
u l
1.005 GAGE LENOTH 2
3 95
\\
i
.25
.256 OfA g DIA
.24G
.256 393 DIA N
R I
e r
r r
r-
,i
~
l a
e I
,e e
e i
- e. '
o s
+
'.19 8
.250 Ry 19 7
.255 TYP L250 REDUCED _~
t.495 1.260 SECTION L480 4.250 4.210
_.630 _
M
.620 ~
O Bt.END UNE FOR R."B" 6
A I.
2 g*
78k i
a f
\\,
.333 As 1-r 4
\\.375 (OF HOLES TO BE DIAf2) FM WITHIN.002 OF 377 IT. 2 ONLY SECTION A-A TRUE ( OF SPECIMEN NOTES:
1-LATHE CENTERS REOUIRED.
2 t2y ALL OVER UNLESS OTHERWISE SPECIFIED.
i FIGURE 4 TENSILE SPECIMEN 4
21
.r A
i 11
< 2-Cf Q:
- fr-
.i 6
l!
L l
P4973 FIGURE 5.
EXTENSOMETER EXTENSION ARr;S AND STRAIN GAGE ASSDiBLY USED FOR TENSILE TESTING
22 Fracture Toughness Tests The WOL specimens were tested according to the recommendations of ASTM E399, "Plane-Strain Fracture Toughness of Metallic Materials".(I')
The procedures for testing the WOL type of fracture specimen are similar to those employed for conventional tensile specimens. The specimen, pulled in tension, is equipped with an extensometer device (COD gage).
The COD gage is attached to two positions on the end of the specimen, such that the two arms span the crack notch. The resultant load-displacement record obtained during the test is then used to confirm that the specimen behaves elastica 1'.y until unstable fracture occurs.
The WOL specimen design is shown in Figure o.
As can be seen in the figure, a notch is machined into the specimen. After final machining and before irradiation, a crack was introduced into the specimen extending from the base of this notch. This precracking was done by f atigue loading the specimen in tension, which normally produces a "sharp crack" condition at the tip of the fatigue crack. The test parameters used to calculate the plain-strain fracture toughness be are 1 ad at unstable fracture, initial crack size, and the specimen dimensions.
The equations employed in this type of K, calculation and the procedural requirements are 7
~
discussed in Reference (19).
An essential part of the testing procedure is the selection of a test temperature that will be likely to permit the occurrence of unstable fracture before gross yielding of the specimen. The latter is essentially centro 11ed by the relationship of K t yield strength. The temperatures Ic required for evaluation of K, values for these irradiated specimens are 7
substantially below room temperature and therefore a cryostat was employed.
This device is used in much the same manner as the furnace is used in elevated-temperature tensile tests. The cryostat was of the split clamshell type; liquid-nitrogen vapor was used to cool the specimen and containment chamber.
The liquid-nitrogen flow rate was adjusted by means of a valve in the gas line.
1 In this manner the nitrogen flow rate and therefore temperature was controlled very accurately during testing.
The temperature drif t during testing was 2F of the test temperature. The COD gage and recorder were calibrated before l
I.45 L43 l.13 0 1.12 0 NOTES:
.380 THRU
' 755 1.005 375 O!A.
.995 I.32/ Al.t.OVER 745 SEE NOTE 2 2 NOTCH DEPTH TO BE EXTENDED
/
BY FATIGUE CRACKING X
{
}
- o 6
.5 01 439
.377 499 +
437+.373 42 i
J-e I.005
-- e-----
,3f7 N
T-7-, E g
o
_]
DG4"X"
'373 l
gg"y* _.
i /
DIA a
,-h R
.5 01
-h DEEP (2)
.500-20 THD. CLASS "B"
.375 DEEP ITEM DIM."X" DIM."Y" 1
.04G3 O23I I
.0473
.0236
.0620
.0 310 2
0630
.0 315 FIGURE 6.
WEDGE OPEN LOADING SPECIMEN
24 testi.'s with a knife-edge micrometer.
All tests were conducted in an Instron run at a crosshead speed of 0.06 in. per min.
The WOL specimen, pull rods, clip gage, and open cryostat are shown in Figure 7.
Figure 8 shows a closeup view of the O L specimen with attached clip gage.
In adCition to the temperature measurements taken during each test, continuous load-time and load-displacement curves were recorded.
4 O
e W
4
l 25 l
l y
i
- i
"if m:
P4879A FIGURE 7.
OPEN CRY 0 STAT SHOWING WOL SPECIMEN WITH ATTACHED CLIP GAGE UPPER PULL R0D
{~
WOL SPECIMEN
? :
Vs.
-Q m
.-CLIP GAGE 6
.g.
I
. ^ CLEVIS l;.
e
.c h,
]
I g
T' LOWER PULL R0D
.I P4879B FIGURE 8.
CLOSEUP VIEW OF WOL SPECIMEN AND CLIP GAGE i
26
_RESULTS AND DISCUSSION Dosimeter and Thermal Monitor Examination The capsule contained th p e S79 F (2.5 percent As, 97.5 percent Pb) and two 590 F (1.75 percent Ag, 0.75 percent Sn, 97.5 percent Pb) thermal monitors. The monitors were in the form of wire with a square or rectangular cross section. The 579 F monitors were located in the top, middle, and bottom regions of the capsule.
The 590 F monitors were located in the top-middle and bottom-middle regions of the capsule.
(The top-middle region is the region located between the top and middle capsule regions, and the bottom. middle region is the region located between the bottom and middle capsule regions.)
The 579 F bottom, the 579 F middle, the 590 F top-middle, and the 590 F bottom -
middle thermal monitors were recovered for examination.
Monitors were examined at a magnification of 4X in a stereomicroscope.
Figures 9 through 12 show the four specimens af ter removal from their Pyrex tubes. The slight bends seen in two of them were introduced when the Pyrex tubes were broken to remove the wires.
None of the 579 F or 590 F thermal monitors show any evidence of melting except the 590 F top-middle monitor. The latter monitor is shown in Figure 12; the right end shows slight evidence of incipient melting.
- However, e
there is no evidence of any general melting along the length of the monitor.
This melting is believed to have occurred during sealing of the temperature monitor in the Pyrex tube.
Based on the examination of the thermal monitor wires, it appears that the capsule was not above 579 F during irradiation for any period of time long enough to cause appreciable melting of the thermal monitors.
Results of the f ast and thermal neutron dosimetry analyses are shown in Tables 2 and 3, respectively.
Of the five iron samples located at 1
different positions, no significant variation in results was noted so that t
an average total neutron fluence (>l Mev) of 3.58 x 10 8,je,2 was selected as most representative of the results.
Excellent agreement was obtained with i
18 37,,2 2he remaining results obtained the nickel dosimeter at 3.57 x 10 from the copper, 238 237 U and Np, ranged from 50 to 100 percent higher. A i
plot of the results of the five dosimeters versus location is shown in Figure 13.
27 4X P4765 4X P4762 FIGURE 9.
579 F BOTTOM THERMAL FIGURE 10.
590 F BOTTOM-MIDDLE MONITOR THERMAL MONITOR l
4X P4764 4X P4767 FIGURE 11.
579 F MIDDLE THERMAL FIGURE 12, 590 F TOP-MIDDLE MONITOR THERMAL MONITOR 4
28 TABLE 2.
FAST NEUIRON DOSIMEIRY RESULTS (>l Mev)
L ***i "
>l Mev nyt >1 Mev I"
2 2
Dosimeter Capsule n/cm -sec n/cm 54 10 18 Fe Top 8.02 x 10 3.77 x 10 54 10 18 Fe Mid Top 7,09 x 10 3.33 x 10 54 0
8 Fe Middle 6.98 x 10 3.28 x 10 0
18 Fe Mid Bottom 8.55 x 10 4.01 x 10 10 10 Fe Bottom 7.52 x 10 3.53 x 10 0
18 Fe Avg 7.63 x 10 Avg 3.58 x 10 58 0
18 Hi Middle 7.60 x 10 3.57 x 10 63 1
18 Cu Top 1.28 x 10 6.03 x 10 63 1
18 Cu Mid Top 1.19 x 10 5.59 x 10 63 0
Cu Mid Bottom 1.28 x 10 6.02 x 10 63 11 18 Cu Bottom 1.30 x110 6.12 x 10 0
11 18 Cu Avg 1.26 x 10 Avg 5.94 x 10 O
18 U
Middle 1.12 x 10 5.27 x 10 237 11 18 pp Middle 1.51 x 10 7.12 x 10 4
4 e
I AP
29 l
TABLE 3.
THERMAL NEUTRON DOSIME111Y RESULTS Location Dosimeter in Capsule
,n/cm -see nyt,n/cm 9
11 19 Co Top 2.68 x 10 1.26 x 10 BARE gg 39 Mid Top 2.70 x 10 1.26 x 10 1
1 Mid Bottom 2.58 x 10 1.21 x 10 '
1 Bottom 2.62 x 10 1.23 x 10 11 19 Avg 2.65 x 10 1.24 x 10 59 0
18 C
Tp 9.47 x 10 4.46 x 10 CADMIUM 0
18 Mid Top 9.65 x 10 4.52 x 10 10 18 Middle 10.6 x-10 5.00 x 10 10 10 Mid Bottom
_9.20 x 10 4.33 x 10 10 1
Avg 9.73 x 10 4.58 x 10 59 11 18 e
1.67 x 10 7.84 x 10 RUE TH M F M *
.
= 2.72 x(
R=C x 0.632 th = C BARE BARE nyt 9
\\
30 1
i i
i o 54 Fe 6 63 Cu O 58 Ni O 2389
'0237g 10 p
Y8 9
?
o e
< 6t R _
f
=
4
.:s 9
.b.
84' D
)
u_
uf2 0
Top Mid Top Mid.
Mid Bottom Bottom 64in
=
Sompte 1.ocationin Capsule FIGURE 13. FAST NEUTRON FLUENCE PROFILES L
I i
-o
31 Several factors enter into the analyses and calculations such as knowledge of the neutron energy spectrum at the surveillance capsule location, effective cross section for each reaction as a function of the neutron energy distribution, purity of certain materials, uncertainty of fission yields, and experimental error.
Therul neutron dosimetry results from the 'Co (n,y) 60Co reaction are given in Table 3, Correcting for the cadmium ratio, R = 2.72, true thermal neutron flux and total fluence values obtained were 1.67 x 1011 18 37c,2, respectively.
A profile of the bare and n/cm -see and 7.84 x 10 cadmium shielded wires in Figure 14 shows very little fluctuation.
Constants used in the six different reactions are summarized in Table 4 Irradiation time of the capsule was 543.6 equivalent full power days and the time of removal from the reactor was September 30, 1972.
The fast neutron cross section values listed are the effective neutron cross sections for neutrons >l Mev.
There is no constant factor applied to the ASTM E261-70 fission spectrum values due to variation in i
threshold energy of the reaction, which is seen to vary from 0.4 Mev to 1.5 Mev and 3-4 Mev for Cu (n,a) 60Co, and the neutron energy distribution at the capsule location.
To calculate the effective cross section, a one-dimensional transport code, ANISN, was utilized.
Twenty-seven energy groups were analyzed.
Plots of the fission spectrum assumed at the core edge and the calculated spectrum at the reactor vessel inner surface are illustrated in Figure 15. Also 54 shown is the Fe (n,p) 54h s
M as a h h d e n m %
ne spectrum at the capsule location is at a lower energy range due to neutron travel i
through about 16 cm of steel and 10 cm of water. The most representative location of the sampics was selected as 2.0 cm outside the thermal shield, his value would represent a conservative figure in that the effective cross section would be slightly smaller (higher resultant flux).
However, a shift of I cm would only affcet the resultant fluence by about 10 percent, o
k
\\
e
32 13X10'8 i
i i
"s 6
1 5900, Bare Wire I
g g l2Xl0
,,i
~
g IE l llXId8 N
10XI0'8 Top Mid Top Mid Mid Bottom Bottom 64in
=
Sample Location in Capsule 7X10'8 i
i i
"e i 6Xl0 ta
.-g EC 18
] 5X10 5900, Cadmium Covered
~
.5 o
O-
-o H
O 18 4XI0 i
i i
Top Mid Top Mid Mid Bottom Bottom h
64in
=!
Sample LocationinCapsule O'
FIGURE 14 TRERMAL NEtTTRON FLUENCE PROFILE i
\\
33 TABLE 4 VALUES USED IN DOSIMETRY CALCULATIONS Target
-(1) Threshold Fission Product Isotope Energy.-
Yield Half-Reaction Target
% Abundance barns Mev Life 03Cu(n.a)60 5.26y Co 100% Cu 69.17 0.00065 Ni(n,p)58co 100% Ni 67.77 0.104 1.0 71.3d Mn Iron (3) 5.82 0.0838 1.5 Fe(n.p)S4 314d Cs U0
>99.9 0.356 0.8 6.30 *.27(2) 30.0y U(n,f)137 38 Np(n,f)
Cs Np0
>99.9 2.63 0.4 6.44 *.315 ) 30.0y 2
9Co(n,y)60Co Al-0.15% Co 100 37.1 5.26y (1) For f ast reactions, 3 is for neutrons >l Mev at capsule location.
(2) Private communication with Wm McElroy, HEDL, June 15, 1973.
(3) Two types of metal were analyzed.
One was SA302 Grade B (97.6 percent Fe) and the other was weld metal (96.6 percent Fe).
4 k
(
34 i
i i
i i
i Sample ANISN
Fission Spectrum O
OReactionCross Section 54fe(n,p) 54Mn s
j QS
- 0.5 Y
~ SW 5 0.4
- 0.4 E s^'l s
I
'5 j 03 -1 Li 03 5 f
a
=
i 1
1 2~
o s
i g
02 1 l
0.2 4.
al
'~l 0'1 L._
I h
' ~ ~ L _.,- - "
O I
0 2
4 6
8 10 12 14 Energy,(Mev) 54 FIGURE 15. NEURON FLUX DISHIBUTION AND Fe CROSS SECTION i
e'
35 237 The value affected the most was that of Np.
In this case the offective cross section is larger than any cross section at an individual energy level because of the "weighting" effect of those cross sections below 1 Hev. The cross sections times the larger flux values below 1 Mev cause theeffectivecrosssection,3,1g, to increast as shown.by the equation ae J
((E) o (E)dE E=0
?>l Mov "
,(E)dE E > 1 Mev where o and c are calculated for each energy group. Only limited data was 03 available on the Cu (n,c) 60Co cross section so that a fission spectrum was assumed. The best value was considered to be calculated from f
[
C (E)dE g
g E=0
,y
>l Mev TOT fa g(E)dE 61 Mev where f = fission spectrum l
f a
= 0.00045 barns from ASIM E261, Table 3.
TOT The ratio of the fission spectra was calculated using the 27 energy groups as follows:
[~
c (E)dE g
61Mev
= 0.688
[
e (E)dE g
E=0 Since the copper results were ~40 percent high, some of the error may occur in (1) uncertainty of the effective cross section and threshold, and/or (2) a small Co impurity.
(It was calculated, using the thermal flux data in this report, that 1 ppm Co in the Cu dosimeter would result in a value
{
see e
36 20 percent high.) Chemical analyses for this material at this low level 237 were not available. The reason for the disagreement of the U and Np 54 values with the g
,g g
significant factors affecting the calculations are the cross section at energies below 1 Mev, dependence on the neutron distribution at the precise sample location,and impurities. Considerable difficulty was encountered in handling the 10-20 mg of U 0 and Np0 Powders but it is thought that the 38 2
experimental error is within 10 percent.
It is doubtful that foreign capsule material contaminated tne samples since this would lower the results.
In su= mary, the iron and nickel dosimeters have the best known physical constants and showed the most consistency resulting in a total fast 8
2 neutron fluence (>l Mev) of 3.58 x 10 n/cm,
O 4
1 e
4 4
6 4
I i
4
37 Impact Tests The results of the tests of the Charpy impact specimens are listed in Tables 5 through 9.
In addition to the impact energy values, the tables also list the measured values of lateral expansion and the estimated fracture-appearance values for each specimen.
These two additional values are sometimes used in measurements of other transition-temperature shifts
- and are included here for the purpose of providing a complete test record for each specimen. The lateral expansion is a measure of the deformation produced by the striking edge of the impact machine when it impacts the specimen; that is, it is a change in thickness measurement of the section directly opposite that where the notch is located.
The fracture appearance value is a visual estimate of the amount of shear or ductile type of fracture appearing on the fracture surf ace of the specimen.
The data listed in Tables 5 through 9 are also graphically shown in Figures 16 through 20.
The shif t in impact properties is clearly shown in these figures. Of primary interes t to this program is the shif t in the temperature corresponding to the impact energy of 30 f t-lb.
This 30 f t-lb fix is that which was established as corresponding to the temperature defined as the nil ductility transition temperature.
During the testing of these Charpy impact specimens, advanced
' instrumentation techniques were used to cbtain valuable supplementary data.
The results of these instrumented Charpy test evaluations are presented in Appendix C-.
The curves for the irradiated pressure vessel material are well defined with little data scatter with only one exception.
That exception is the curve for the irradiated HAZ metal.
The reason for this is that it is difficult to cut specimens out of a heat-affected zone in a plate between base metal and weld metal, and be assured that the HAZ specimens all have the identical microstructure and thermal history. The main purpose of the Charpy impact tests was to determine the NDTT of the two base metal plates, the HAZ i
l metal, and the veld metal of the pressure vessel. The results of these NDTT l
evaluations are summarized in Table 10.
Al'so listed in the table are the results of the impact tests on the AS1H correlation monitor material.
I 2
38 TABLE 5.
CHARPY V-NOTCH IMPACT TEST RESULTS FOR BASE METAL PLATE A9811 Test Impact Lateral Fracture Specimen Temperature,
- Energy, Expansion, Appearance, Fracture Specimen F
ft ib mils Percent Shear Surface r:a';;,....
A2
-65 4.5 5
0 j,
y,,,,
. f'5-A5
-25 5.5 5
2
'f
.i Pi,
A7 5
8 8
15
.-1
./ :[ I, / 'i.
A3 26 22 21 15 i,e-e n/
7 '-
A9 33 26 26 25 l
q c.:l' <, j
~
m A6 51 41 38 35
[ f,
),4 w
r,~4
,=j_
All 79 46 45 40 C...
7
39 TABLE 5 (CONTINUED) l Test Impact Lateral Fracture Specimen i
Temperature,
- Energy, Expansion, Appearance, Fracture Specimen F
ft Ib mils Percent Shear Surface A12 104 55 51 70 I /,'
\\
l
\\
f A10 119 71 64 70 ggy g
.5 f
---4
40 h
TABLE 6.
CHARPY V-NOTCH DiPACT TEST RESlTLTS FOR BASE METAL PLATE C1423 Test Impact Lateral Fracture Specimen Temperature,
- Energy, Expansion, Appearance, Fracture Specimen F
ft Ib mils Percent Shear Surface Y,k ,$*[
C3
-102 3
4 0
<3 k,,,. (.,
37, C1
-65 3
2 2
ig
,T f.d d.??. lYi
\\
f
- }s l
C10
-24 11.5 10 10
,'.6 [>.'J(s
{
l
, -f.
{ 4 *.9. >.
hh t.?, '
4.
i C5 0
18 16 10 U
V.4 i > :. t i qa-l.
3 --
.I Jg >'.r..'
C7 5
37 31 15
' M '- l [-
f Yb.
h*
~-
q>
C12 26 35 30 15 e... I I
C11 33 38 31 40
]
41 i
)
TABLE 6 (CONTINUED)
Test Impact Lateral Fracture Specimen Temperature,
- Energy, Expansion, Appearance, Fracture Specimen F
f t Ib mils Percent Shear Surface C8 52 55 45 25 14{.;*,],,,,,
Vci:
s t
?! ) f?
C2 79 76 63 55
- - }
i 1
C4 120 101 73 80 gg gg 7.
i i
i
- - - ~ -
- - - - - ~ '- - - -
- ~~
42 TABLE 7.
CHARPY V-NOTCH IMPACT TEST RESULTS FOR HEAT-AFFECTED ZONE METAL Test Impact Lateral Fracture Specimen j
Temperature,
- Energy, Expansion, Appearance, Fracture Specimen F
ft Ib mils Percent Shear Surface "j
W4
-87 5.5 3
5
.- Ty.
c.4,r i
s w.... _
"'C f:;p 'P
- c,
l[-
,j' r.
WS
-26 31 28 25
.E ' tL.
. __ y t. (>I W7 0
33 31 40 J
_ t_
,r 7-,_-[)?
x WS 25 20 15 35 J%,4.
M W3 36 99 70 90 j
W6 51 83 50 80
.i W1 79 110 81 95 W2 208 105 81 98 l
43 TABLE 8 CHARPY V-NOTCH IMPACT TEST RESULTS FOR k' ELD METAL Test Impact Lateral Frac tu re Specimen Temperature,
- Energy, Expansion, Appearance, Fracture Specirnen F
ft Ib mils Percent Shear Surface W8
-87 2.5 1
5 W7
-31 6.5 5
15 W2 25 13 12 25 W4 50 27 22 40 M
W6 79 34 31 50 M,
l WS 123 46 43 70 l
Mi i
l W1 167 54 53 98 i
l k
i I
W3 208 52 53 99 i
64 TABLE 9 CHARPY V-NOICH IMPACT TEST RESULTS FOR ASIM CORRELATION HONITOR MATERIAL Test Impact Lateral F cture Specimen Temperature,
- Energy, Expansion, Appearance, Fracture Specimen F
ft Ib mile Percent Shear Surface
' t if, ~,'<,.. )
{
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e 26
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R4 127 27 26 40
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- i R2 161 40 36 55 gg R.
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100 $
Unirrodioted
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80.-
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- 10 0 b
0 2
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FIGURE 16. CllARPY IttPACT TEST RESULTS FOR BASE-t!ETAL PLATE A9811 i
~
t l
l
46 I
i i
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--- Unirradicted
,,10 0.8
, s
- -o m
80 88 Irradiated
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= ANDTT 4
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-200
-10 0 0
10 0 200 300 Ternperature, F FIGURE 17. CilARPY DlPACT TEST RESULTS FOR BASE-METAL PLATE C1423
47 s
e
- --- - Unirradiated
'/
10 0 *8 3
i o
Irradiated 80 -
- /
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M oNDTT o
go i
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10 0 200 300 Temperature, F FIGURE 18. CilARPY IMPACT TEST RESULTS FOR llEAT-AFFECTED ZONE HETAL We 4
y
9 48 i
i i
is
---~ Unirradiated 100 E
/,s
-m--
m Irradiated
/
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10 0 0 C
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60 p
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/ANDTT
~
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s'/
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- l 0
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10 0 200 300 Temperofure, F FIGURE 19. CRARPY IMPACT TEST RESULTS FOR WELD HETAL l
l
1 49 i
i i
i Unirradiated 100.$
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20 o
n 0
10 0
~
80.L
/
60 $
/
3a i
40 w
/ A NDTT go h
/
o o
~
-200
-10 0 0
10 0 200 300 Temperature, F FIGURE 20. CHARPY IMPACT TEST RESULTS FOR CORRELATION-MONITOR MATERIAL J
D 4
e es J
50 l
l TABLE la
SUMMARY
OF CHARPY IMPACT TEST RESULTS 1
Unitradiated Irradiated
Plate A9811
-45 45
+90 Plate C1423
-30
+20
+50 HAZ Metal
-60
+10
+70 Weld Metal
-45
%5
+110 Correlation Monitor 40
+135
+95 h
J.
'I a
l 1
l i
4 e
'4 1
B moo
-... i
1 51 Irradiation caused the NDTT to increase in all four of the pressure vessel materials. The base metal plate A9811 had an NDTT increase of 90 F, the base metal plate C1423 had an NDTT increase of 50 F, and the HAZ metal had an NDTT increase of 70 F.
The weld metal had an NDIT increase of 110 F, the greatest of the four pressure vessel materials.
The highest NDTT of the unirradiated pressure vessel materials was -30 F for the base metal plate C1423. The highest NDTT of the irradiated pressure vessel materials is +65 F for the weld metal. Therefore, the limiting metal of the four pressure vessel materials at the present irradiation level is the veld metal.
The upper shelf energy of a Charpy impact curve is defined as the upper level of energy that curves exhibit at higher temperatures where increases in test temperature cause no further increase in impact energy.
A general effect of irradiation is to lower the upper shelf energy. There is a sufficient amount of scatter in the actual data points for the preirradiation and postirradiation data to mr.ke the confirmation of this trend difficult in the present investigation in the case of the HAZ metal and the base metal plate C1423.
(Data points for the preirradiation Charpy curves are not shown in Figures 16 through 20.)
However, the base metal plate A9811 and the weld metal both show clear decreases in the upper shelf energy.
The upper shelf energy for all four irradiated materials is above 50 ft ib. However, the upper shelf energy for the weld metal is only slightly above, being about 53 ft ib. As discussed earlier, some investigators have stated that the NDTT concepts no longer apply when the upper shelf energy drops below 50 ft lb. When the impact specimens in the next surveillance capsule are tested, the test temperatures should be chosen such that the upper shelf energy is well established to determine if it has dropped below the 50 f t-lb level.
The capsule also contained ASTM correlation monitor Charpy impact spec imens. The unirradiated and irradiated test results are shown plotted in Figure 20. As indicated in Table 10, the preirradiation NDTT was +40 F and the I
postirradiation NDTT is +135 F, resulting in a nil ductility transition temperature shif t of +95 F.
-~
.~.
i 52 1
1 3
The 6NDTT values for the irradiated materials other than the ASTM correlation monitor material are shown plotted in Figure 21 as a function i
of fluence. Also shown in this figure are ANDTT values obtained in other surveillance prograris(29-35)
The apparent large scatter in data among the l
various programs is not unusual. Note that the weld metal values determine the upper bound of the trend band. The values used to form the trend band
]
are'those from programs where the irradiation temperature was between 550 and 590 F.
It can be seen that the four 6NDTT values for the pressure vessel I
materials of the present program fall well within the upper and lovar bounds determined by materials of other investigations.
The 6NDTT value for the irradiated correlation monitor material is shown in Figure 22. Also shown in this figure are 6NDTT values obtained 3
from other programs using this material.
The ANDTT value determined in the i
present program is in good agreement with the other values shown.
The Figure 21 pressure vessel trend band for,6NDTT charges can be j
l
)
used to estimate the 6NDTT value of the pressure-vessel after various periods t
of reactor operation. These estimates assume that the surveillance capsule f
receives an accelerated fluence of 3.3 times that of the pressure vessel vall(36)
)
l l
and that the irradiation temperature of the pressure-vessel vall will continue
[
to be equal to or greater than 550 F.
The 6NDTI estimates for 10, 20, and 32
?
]
, years of effective full-power operation of the reactor are shown in Figure 23.
The highest estimated values for the materials are reprssented by the upper
}
bound of the trend band. The predicted,4NDTT values for 10, 20, and 32 years of f
j effective full power operation [1520 Mw(t)) are 210, 240, and 255 F.
This is in reasonable agreement with the 32 years preirradiation prediction of 280 F for e
the reactor.00) Subsequent examinations of specimens from future capsules will be used to substantiate the present 32 years NDIT shif t estimate.
l i
i i
i e
t 4
i
+
53 l
i i
l I I iiIl l
l l
6 4 l l1l Base Weld HAZ Point Beach No.1 Q
Q 90nnecHeut Yank.ee o
o e
300 Big Rock Point 7
V V
Yankee Surveillance 6
Yankee Special O
Humbolt Bay b
San Onofre 4
4 y
g.
Elk River X
i200 E
V 8
I
.4 V
ke X
b A
z
-10 0
.Q e
Q V
X x
Trend Band for 550-590 F l
x j
I I IIil I
I I
I IIiil 0
i K)18 jols 1020 Neutron Fluence, nyt i
FIGURE 21. COMPARISON OF ANDTT VALUES FROM VARIOUS SURVEILLANCE PROGRMIS FOR SA302 GRADE B PRESSURE-VESSEL MATERIALS
54 I
I I
i iei1l 1
1 I
I iii1[
Q - Point Be0ch No.1 o
Connecticut Yankee V - Big Rock Point 300 O - Yankee
< - Son Onofre
"- 200 S
3 C
e E.
e-
~
$5 z 10 0 O
o 7
Trend Bond For 550-590 F V
0 1
I I Ilill i
I I
I litil 1018 109 1020 1
Neutron Fluence,nyt J
IIGURE 22. COMPARISON OF 6HDTT VALUES FROH VARIOUS SURVEILLANCE PROGRAMS FOR SA302 CRADE B ASTM CORRELATION-MONITOR HATERIAL w-
I 55 i
i i
i iiii i
i i iiiit 300 Equivalent Full Power Years Of Operation 32 g
- 10 l
=
200 xxl r
f NNN N,
I 10 0 -
l Trend Bond For l 550-590 F l
l I
l i
!iii i
i i
in,iL i 0
i8 10 10'8 10 2
Neutron Fluence,nyt FIGURE 23. 6NDTT ESTIMATES FOR 10, 20, AND 32 YEAR OF FULL-POWER OPERATION j
_. The predicted maximum ANDIT values for 10, 20, and 32 years of effective full power operation are 210, 240, and 255 F, re spec tively.
l
~.
T L
56 t
+
Tensile Tests 3
i 1
[
The capsule contained tensile specimens from the base estal plate f
A9811, the base metal plate C1423, and the weld metal.
Specimens were tested at temperatures ranging from -180 F to 550 F.
The tensile data from these tests are listed in Table 11.
The stress-strain curves are
\\
shown in Figures 24 through 30. In addition, photographs of the fractured i
specimens are given in Figures 31 through 37.
These photographs show the t
j necked down region of the gage length and the fractures.
t j
Figures 38 through 40 show the unirradiated and irradiated _ values
]
of elongation, reduction in area, 0.2 percent yield strength, and ultimate l
tensile strength for the three materials.
The unieradiated values are from the report WCAP-7513.(9)
The changes in tensile properties in general follow the expected behavior of irradiated metals such as the pressure vessel steel j
l SA302 Grade B in the range 75 F to 550 F.
There is no preirradiation low temperature data to which the -110 and -180 F irradiated tensile results i
4 can be compared. The elongation and reduction in area values generally have l
r decreased as a result of irradiation, with the exception of plate C1423 in which the reduction in area at 75 F is unchanged and at 550 F is slightly increased.
l
.} ~
The 0.2 percent off set yield strength and ultimate tensile strength have j
r increased in all cases.
There is some scatter in the reduction in area values l
' for base metal plate A9811, in that the 75 F value is mederately greater than f
)
the -110 and 550 F values, i
I i
j r
i
}
i i
i
(
't Q
l b
I TABLE 11.
TENSILE PROPERTIES OF IRRADIATED PRESSURE-VESSEL SPECIMENS Test 0.2 Percent Ultimate Total Reduction Temperature, Offset Yield Tensile Elongation, in Area, Material Specimen F
Strength, psi Strength, psi Percent Percent Plate A9811 A3
-180 91,900 112,000 25.8 56.1 Plate A9811 A2
-110 79,800 102,700 24.1 48.6 13 Plate A9811 A4 75 67,800 89,100 18.9 58.5 Plate A9811 A1 550 59,500 88,400 11.9 44.6 Plate C1423 C2 75 76,000 96,200 17.0 65.0
- Plate C1423 C1 550 66,800 93,800 13.1 59.0 Weld Metal WW1 550 81,900 100,800 12.4 51.0 m
9 9
,--,m-
58 200,000 m.
100,000
[
m
~
- I IG f
0 0
5 10 15 20 25 30 Percent Elongation,in/in 1
FIGURE 24 STRESS-S1 RAIN CURVE FOR IRRADIATED BASE-METAL PLATE A9811 (SPECIMEN A3) TESTED AT -180 F 9
~
'I k 1,.
i r
60 E
200,000 i
a vr a 100,000 -
.:=
v>
1 g
0 I
0 5
10 15 20 25 30 l
Percent Elongation, in/in l
FIGURE 26. SBESS-SRAIN CURVE FOR IRRADIATED BASE HETAL PLATE A9811 (SPECIMEN A4) TESTED AT 75 F
- i i
I e
e a
f i
e oom s
j I
61 200,000 i
i i
1 1
.g a
VI E 100,000 -
x v2 1
0-I I
J O
5 10 15 20 25 30 Percent Elongation,in/in FIGURE 27.
STRESS-STRAIN CURVE FOR IRRADIATED BASE-METAL PLATE A9811 (SPECIMEN A1) TESTED AT 550 F l
l
)
I
., f
i i
63 200,000' 3
i i
i i
E.
6 100,000
=
m l
l 0
O 5
10 15 20 25 30 Percent Elongation, in/in 9
FIGURE 29. STRESS-SIRAIN CURVE FOR IRRADIATED BASE-METAL PLATE C1423 1
(SPECIMEN C1) TESTED AT 550 F 4
4 a
f J
_--e
64 4
200,000=
~
i i
u m.
[ 100,000 N
4
- )
l l
I 1
l l
0 5
10 15 20 25 30 Percent Elongation,in/in 1
l FIGURE 30. STRESS-STRAIN CURVE FOR WELD METAL (SPECDiDI W1) i TESTED AT 550 F i
- i
l 65 i
1
.(.
\\
l j
i a
FIGURE 31.
BASE METAL TENSILE SPECIMEN A3 TESTED AT -180 F 1
1 1
4 i
i l
.i 1
i P4900 j
FIGl'RE 32 BASE METAL TENSILE SPECIMIN A2 TESTED AT -110 F I
l i
l l
I 1
_m.
- - =__
_-mm j
66
- l l
l I
i i
i k
I gs t
1 i
i i
P4805 i
FIGt7E 33. BASE METAL TENSILE SPECIMEN A4 TESTED AT 75 F i
E a
i t
l c
l 4
i l
i
(
e P4804 i'
j FIGl1E 34 BASE METAL TENSILE SPECIMEN Al TESTED AT 550 F l
l
a.h,a_-2
.wa.a.wa--m ar-a.4-ummamm.-.A-a ma a4 4m 4.m a w mm.m.
m.-4
-.me%.am a..
A__
--.__m..h_.
.a
-_h+.._h-a_ami 1
I j
P4807 l
FIGURE 35.
BASE METAL TENSILE SPECIMEN C2 TESTED AT 75 F 4
J F4806 FIGURE 36 BASE METAL TENSILE SPECIMEN Cl TESTED AT 550 F l
1 t
j i
l, _ _... _ _ _ _ _ _
.l f
}
68 i
I i
}
l I.
9 i
l i
i i
l I
e l
P4809 FIGURE 37 kTLD METAL TENSILE SPECIMEN W1 TESTED AT 550 F 1
l l
l l
?
'I
-~
_. - ~, ~ - ~ _.
69 120p00 :
i i
i l
i i
i 100000 K
4 A
80
- g
,000 e
ULTit[ ATE TENSILE STR 6
5 60p00 -
M d
~
3 0.2% YlELD STRENGTH 40p00 Unitradiated-o a V O Irrodiatod-0AVO c:
S
(
80
~
o V
o n
v V
Q v
60 y
l q
REDUCTION Ill AREA
~
g V
q j40
%N-T
-O 3
D 20 C
~
ELONGATION O
0-
-200
-10 0 0
10 0 200 300 400 500 600 Temperature, F FICURE 38.
UNIRRADIATED AND IRRADIATED TENSILE PROPDITIES FOR THE VEP POINT BEACH UNIT NO, 1 REACTOR PRESSURE VESSEL SHELL P1/TE A9811 s
1 4
1
70 100,000 i
i i
i i
j c
s ULTIMATE TENSILE STRENGTH 80,000 b
a f60,000 -
U 0.2% YlELD STRENGTH l
i 40,000
~
~
Unirradioted-o A V O Irradioted-oAV O 80 y
60
^7 REDUCTION IN AREA E8g 40 2N-
_g 20 ELONGATION g_
O I
l-I I
)
00 10 0 200 300 400 500 600 Temperature, F FIGURE 39 UNIRRADIATED AND IRRADIATED TDiSILE FROPERTIES FOR Tile kTP POINT BEACll UNIT NO.1 REACTOR PRESSURE VESSEL SilELL PLATE C1423 e
71 120,000 i
i i
i i
100,000 Q
ULTIMATE TENSILE STRENGTH c; 80,000 v
E 0.2% YlELD STREf!GTH O
8-60,000 40,000 Unirrodiolad-o a V O
~
Irradiated-oAVO g
g 80 REDUCTION IN AREA 60
?,
V E
ct 40 B
n ELONGATION D
20 f
Q)
G I
I I
I I
I
~
0 10 0 200 300 400 500 600 Temperature, F 1
TICL'RE 40.
- ~dIERADIATED AND IRRADIATED TENSILE FROPIRTIES FOR THE WEP v
5-FolNT 11DCll UNIT NO 1 REACTOR FRESSURE VESSEL VELD HETAL
72 Fracture Toughness Tests A typical WOL load-deflection curve is shown in Figure 41.
The curve shown is for specimen C1 which was tested at -149 F.
The curve meets the AS E requirements for a valid fracture toughness test.
If substantial plastic deformation had occurred before f racture, the plot would be curved rather than a straight line. The testing of the nine. irradiated WOL specimens resulted in the determination of nine valid K fracture toughness g
values. The term "valid" refers to compliance with the Asm E399-70T recommendation (I9) the K
B 2 2.5 (,I*)
ys where B is specimen thickness (1 in. for these specimens) and c, is the y
tensile yield strength. The results of these WOL specimen tests are listed in Table 12.
The crack length used in the K, calculation was measured on the 7
fracture surface after the test was completed.
Three measurements were taken and averaged to determine the average crack length of each specimen. Views of the fracture surf aces of the irradiated specimens are shown in Figures 42 through 50. The curved line across each specimen half is the front of the
. crack introduced before irradiation by fatigue precracking.
The equation used to calculate K, is 7
K,=
f(a/W) 7 where a = the original crack length, from the load centerline to the average extent of the fatigue crack B = specimen thickness P = load at fracture.
W = specimen depth, from load centerline to specimen end f(a/W) = dimensionless stress-intensity f actor that erapirically relates the applied load to the stress state at the
' fatigue crack.
o mm I
l
73 5000
=-
i i
i i
i Specimen! C-l Temperature:-149F 4000 l
l 3000 u,
- f3
.E
.3 2000 1000 l
0-0
.002
.004
.006
.008
.010
.012 Clip Gage Deflection, inches FIGURE 41 TYPICAL LOAD-DEFLECTION CURVE OBTAINED DURING A WOL TEST O
4 e
74 TABLE 12. RESULTS OF WOL SPECINEN TESTS OF BASE METAL AND WELD METAL Test Average Crack Fracture g
Temperature, Length (a),
Depth (W)
Load (?),
Ic Specimen F
in.
in.
Ib psi [lii.
A3
-180 0.543.
1.135 3620 30,600 A1
-150 0.533 1.130 4690 39,100 A4
-150 0.540 1.128 4140 35,200 A2
-110 0.571 1.154 4680 41,200 C1
-149 0.546 1.129 4150 35,800 C3
-130 0.561 1.164 3980 33,600 C2
-108 0.536 1.136 5380 44,700 W1
-90 0.563 1.150 3970 34,500 W2
-45 0.549 1.126 4840 42,300 9
9 e
9 s
s-r
75
~'-- --
en i
a ei.
P4886 P4889 FIGURE 42 BASE METAL WOL SPECIMEN FIGURE 43. BASE METAL WOL SPECIMEN A3 TESTED AT -180 F Al TESTED AT -150 F
-4
~
Y ;.8 W r:-
h[4'hy
- ~
- ~n
,. m.u,_
aq P4882 P4890 FIGURE 44 BASE METAL WOL SPECIMEN FIGURE 45 BASE METAL WOL SPECIMEN A4 TESTED AT -150 F A2 TESTED AT -110 F
76 l
ms. -,
\\
_ u-i-
,.s P4888 P4883 FIGURE 46 BASE METAL WOL FIGURE 47.
BASE METAL WOL SPECIMEN C1 TESTED SPECIMEN C3 TESTED AT -149 F AT - 130 F C
I
./ -
1 P4884 FIGURE 48 BASE METAL WOL SPECIMEN C2 TESTED AT -108 F
--,-,,,.,a
---n,
-w n e m,n n.-
p g g y,,mm
l i
i
)
)
77 1
i l
i l
M M.
, ;c[
' ' hd.f,.
.. a k. h _ _,_ _
'13L i I
i P4887 l
FIGURE 49 WELD METAL WOL SPECIMEN W1 TESTED AT -90 F l
i
/
e em 4%
P4885 s
FIGURE SR WELD METAL SPECIMEN WW2 TESTED AT -45 F l
L
78 The selection of test temperatures was based on the results of the impact specimen tests.
In the surveilitnce program outlined in WCAP-7513 it is recommended that the first WOL specimen of each material be tested at a temperature of -200 F plus the nil ductility transition temperature shift that was measured by the impact specimen tests. For the A-series base metal the recommended temperature was therefore -110 F
(-200 + 90); for the C series. -150 F; and for the weld metal series -90 F.
The recommended temperatures were used as the initial test for each series.
The results of the WOL specimen tests of the three irradiated materials are shown in Figure 51 in the form of K versus temperature.
The Ic preirradiation values are not shown, as they are not reported in WCAP 7513.
The fracture toughness in general increases as temperatures increases for each of the three materials.
The fracture toughness values for the two base plate materials are in the same general range for a particular temperature, with the weld metal values lower than the two base plate materials.
As can be seen in the figure, there is some scatter in the data.
This is shown by the two tests for base metal plate A9811 at -150 F, which yielded fracture toughness values of 40,000 and 44,500 psi /in.
This amount of scatter in the fracture toughness values for this type of pressure vessel steel is not unusual. However, it indicates that caution should be used in making conclusions based on a limited number of tests as in the case of the present program.
I 69a
,m.
79 60,000 i
i o
O O
O 6
40,000 g
O C3 a
~u M
20,000 o
Base meto' plate A98ll A
Base metal plate Cl423 a
Weld metal 0-
-200
-16 0
-12 0
-80
-40 0
Temperature, F FIGURE 51. FRACTURE TOUCilNESS OF T11REE IRRADIATED MATERIALS AS A FUNCTION OF TDIPERATURE
\\
80
_ CONCLUSIONS The maximum irradiation temperature of the surveillance capsule during irradiation did not exceed 590 F.
The neutron fluence experienced 18 by the capsule was 3.58 x 10 nyt (>l Mev) which was attained after 1.49 years equivalent full-power years of operation.
This is equivalent to a maximum fluence of 1.11 x 10 nyt (>l Mev) for the pressure vessel af ter 1.49 equivalent full power years. For a pressure-vessel life of 40 years operation at an 80 percent load factor (32 equivalent full-power years),
the maximum fluence experienced by the pressure vessel would therefore be 19 predicted to be 2.38 x 10 nyt (>l Mev).
Tensile tests were conducted on the two base metal plates and the veld metal. The yield strength and ultimate tensile strength of all three materials increased at 550 F due to irradiation.
The corresponding ductilities (reduction in area and total elongation) decreased in all three materials at 550 F, except for the reduction in area of base metal plate C1423 which increased slightly.
Fracture toughness tests were conducted on the two base metal plates and the veld metal in the general temperature range of -180 F to -45 F.
The plane strain fracture toughness for these materials was found to be in the general range of 30,600 to 44,700 psi [1n., with the fracture toughness generafly increasing with' increasing temperature.
The shif t in the nil ductility transition temperature (NDTI) was determined for the two base metal plates, heat-affected zone metal, and weld metal, ne greatest NDTT shift due to irradiation was found to occur in the weld metal, which had a shift of +110 F.
The highest NDTT af ter irradiation was the +65 F value determined for the weld metal.
The NDTT shifts observed for the pressure vessel materials in the present investigation fit within a trend band for NDTT values from similar surveillance programs of other reactors using the same pressure-vessel material, SA302 Grade B.
The results obtained in this investigation show that the pressure vessel base metal, heat-affected zone metal, and weld metal mechanical properties are changing with irradiation in a manner consistent with the changes expected when the pressure vessel was fabricated and put into service.
e--
7
81 REFERENCES (1) Reuther, T.
C., and Zwilsky, K. M., "The Ef fects of Neutron Irradiation on the Toughness and Ductility of Steels", in Proceedings of Toward Improved Ductility and Toughness Symposium, published by Iron and Steel Institute of Japan (October, 1971) pp 289-319.
(2) Steele, L. E., "Major Factors Affecting'deutron Irradiation Embrittlement of Pressure-Vessel Steels and Weldments, NRL Report 7176 (October 30, 1970).
(3) Berggren, R. G., "Critical Factors in the Interpretation of Radiation Effects on the Mechen'. cal Propercies of Structural Metals", Welding Research Council Bul'.etin 8Z, 1 (1963).
(4) Witt, F.
J., "Heavv-Section Steel Technology Program Semiannual Progress Report for Period inding February 29, 1972", ORNL Report 4816 (October,1972).
(5) Alger, J. 'V., and Skupien, L. M., "Neutron Radiation Embrittlement at 500 and 650 F of Reactor Pressure Vessel Steels", in Materials in Nuclear Applications, American Society for Testing and Materials Special Technical Publication 276 (1960), pp 116-134.
(6) Porter, L.
F., "Radiation Effects in Steel", in Materials in Nuclear Applications, American Society for Testing and Materials Special Technical Publication 276 (1960), pp 147-196.
(7) Steele, L. E., and Hawthorne, J. R., "New Information on Neutron Embrittle-ment and Embrittlement Relief of Reactor Pressure Vessel Steels", American Society for Testing and Materials,Special Technical Publication 381 (1964).
(8) ASIM Designation E185-70, "Surveillance Tests on Structural Materials in Nuclear Reactors", Book of ASn! Standards, Part 31 (1972), pp 565-569.
(9) Yanichko, S. E., "Wisconsin Michigan Power Co. Point Beach Unit No.1 Reactor Vessel Radiation Surveillance Program, WCAP 7513 (June,1970).
(10) Potapovs, U., and Hawthorne, J. R., "The Effect of Residual Elements on 550 F Irradiation Response of Selected Pressure Vessel Steels and Weldments",
i Nuci Appi,,6, 27-46 (1969).
(11) Hawthorne, J. R., and Fortner, E., "Radiation and Temper Embrittlement Processes in Advanced Reactor Weld Metals", J. Eng. Indus., Trans. ASME, 94, 807-814 (1972).
(12) Serpan, C.
Z., et al., "Radiation Damage Surveillance of Power Reactor Pressure Vessels", NRL Report 6349 (1966).
(13) Barton, J.
R., and Wylie, R. D., "Development of Nondestructive Testing Instrumentation for Reactor Pressure Vessels", SwRI-1243-14 (1964).
8--
82 (14) Steele, L. E., et al., "Neutron Irradiation Embrittlement of Several Higher Strength Steels", NRL Report 6419 '(September 7,1966).
(15) ASE Designation E208-69, "Conducting Drop-Weight Test to' Determine Nil Ductility Transition Temperature of Ferritic Steels". Book of ASTM Standards, Part 31 (1972), pp 594-613.
(16) Pellini, W. S., and Puzak, P. P., "Frac ture Analysis Diagram Procedures for the Fracture-Safe Engineering Design of Steel Structures",
Bulletin 88, Welding Research Council (May, 1963),
(17) Loss, F. J., Hawthorne, J. R., and Serpan, C. Z., Jr., "A Reassessment of Fracture-Safe Operating Criteria for Reactor Vessel Steels Based on Charpy-V Performance", NRL Report 7152 (September 8, 1970).
(18) Sailors, R. H., and Corten, H. T., "Relationship Between Material Fracture Toughness Using Fracture Mechanics and Transition Temperature Tests",
in Fracture Toughness, AS m Special Technical Publication 514 (1972) pp 164-191.
(19) ASE Designation E399-70T, "Plane-Strain Fracture Toughness of Metallic Materials", Book of ASTM Standards, Part 31 (1970), pp 911-927.
(20) ASIM Designation E320-69T, "Radiochemical Determination of Cesium-137 in Nuclear Fuel Solutions", Book of ASm Standards, Part 30 (1970), pp 1004-1009.
(21) ASE Designation E261-70, "Measuring Neutron F. lux by Radioactivation Techniques", Book of ASE Standards, Part 30 (1970), pp 762-772.
(22) ASE Designation E262-70, "Measuring Thermal-Neutron Flux by Radioactivation Techniques", Book of.4S M Standarde, Part 30 (1970), pp 773-780.
(23) ASE Designation E263-70, "Measuring Fast-Neutron Flux by Radioactivation of Iron", Book of ASE Standards, Part 30 (1970), pp 781-783 (24) ASm Designation E264-70, "Measuring Fast-Neutron Flux by Radioactivation of Nickel", Book of ASM Standards, Part 30 (1970), pp 787-791.
(25) ASE Designation E343-07T, "Fast-Neutron Flux by Activation of Molybdenum-99 Activity from Uranium-238 Fission", Book of ASm Standards, Part 30 (1970),
pp 1078-1084.-
(26) ASTM Designation E393-69T, "Measuring Fast-Neutron Flux for Analysis for i
Barium-140 Produced by Uranium-238 Fission", Book of ASM Standards, Part 30 (1970), pp 1174-1180.
t (27) ASTM Designation E23-66, "Notched Bar Impact Testing of Metallic Materials",
l Book of ASIM Standards, Part 31 (1970), pp 271-285.
(28) ASTM Designation A370-71, "Mechanical Testing of Steel Products", Book of ASTM Standards, Part 31 (1971), p 45.
r-'7"
-' "*' * + " * " * - - * " ~ - -
d
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*C
- -7~
1 l
\\
83 I
(29) Serpan, C. Z., Jr., and Watson, H. E., "Mechanical Property and Neutron Spectral Analyses of the Big Rock Point Reactor Pressure Vessel",
Nucl. Eng. Design, 11, 393-415 (1970).
(30) Serpan, C. Z., Jr., and Hawthorne, J. R., "Yankee Reactor Pressure-Vessel Surveillance: Notch Ductility Performance of Vessel Steel and Maximum Service Fluence Determined from Exposure During Cores II, III, and IV", NRL Report 6616 (September 29, 1967).
(31) Brandt, F.
A., "Humboldt Bay Power Plant Unit No. 3 Reactor Vessel Steel Surveillance Program", GECR-5492 (May,1967).
(32) Ireland, D. R., and Norris, E. B., "Influence of Neutron Irradiation on the Properties of Steels and Weld Typical of the ERR Pressure Vessel Af ter Two Power Years Operation", SwRI-1228-P-9-15 (March,1968).
(33) Sterne, R. H., Jr., and Steele, L. E., "Steels for Commercial Nuclear Power Reactor Pressure Vessels", Nucl. Eng. Design, 10, 259-307 (1969).
(34) "Analysis of First Surveillance Material Capsule from San Onofre Unit I",
Southern California Edison Company (July,1971).
(35) Perrin, J. S., Sheckherd, J. W., and Scotti, V. G., "Examination and Evaluation of Capsule F for the Connecticut Yankee Reactor Pressure-Vessel Surveillance Progrcm", Final Report to Connecticut Yankee Atomic Power company (March 30, 1972).
(36) Private Communication from J. J. Zach and G. A. Reed, Point Beach Nuclear Power Plant (March 8, 1973).
l 1
l N
9 3
l APPENDIX A PRESSURE VESSEL MATERIAL O
e 4
4 4
ee e
APPENDIX A*
PRESSURE VESSEL MATERIAL Babcock and Wilcox Co.
supplied sections of SA302 Grade B steel plate to Westinghouse for the reactor. vessel surveillance program. These sections represented material from the 6-3/4-inch-thick intermediate shell course plate A9811 and the 6-3/4-inch-thick lower shell course plate C1423 used in the WEP pressure vessel.
In addition, a weldment made from sections of the two plates was also supplied by Babcock and Wilcox Co.
The plates were produced for Babcock and Wilcox Co.
by the Lukens Steel Co.
The chemical analyses and heat treatment history of the plate material follows.
- a. Chemical Analyses Percent Plate Plate Weld Element A9811 C1423 Metal C
0.19 0.21 0.09 Mn 1.42 1.37 1,47 1
P 0.010 0.014 0.019 S
0.020 0.019 0.024 Si 0.25 0.25 0.49 Mo 0.48 0.46 0.39 Cu 0.18 Ni 0.57 Cr 0.13 A1 0.035 N
0.016 2
V 0.001 Sn 0.004 Ti 0.001 As 0.004
- The information in this Appendix is from Yanichko, S.E., "Wisconsin Michigan Power Co. Point Beach Unit No.1 Reactor Vessel Radiation Surveillanco Program", WCAP-7513, Westinghouse Electric Corporation (June, 1970).
4
A-2 Co 0.001*
Zr 0.001*
Sb 0.001*
Zn 0.001*
B 0.003*
- Not detected. The number indicates the minimum limit of detection,
- b. Heat Treatment Plate A9811 Heated at 1650 F, 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br />, water-quenched Tempered at 1225 F, 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br />, aircooled Stress-relieved at 1125 F,11-1/4 hours, furnace-cooled Plate C1423 Heated at 1650 F, 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br />, water-quenched Tempered at 1225 F, 7 hours8.101852e-5 days <br />0.00194 hours <br />1.157407e-5 weeks <br />2.6635e-6 months <br />, aircooled Stress-relieved at 1125 F,10-1/2 hours, furnace-cooled Weldment Stress-relieved at 1125 F, 11-1/4 hours, furnace-cooled h
W*
t 0
a
APPENDIX B CORRELATION MONITOR MATERIAL 4
i i
. - 1
l APPENDIX B*
CORRELATION HONITOR MATERIAL The correlation monitor material SA302 Grade B was furnished by the U. S. Steel Corporation through Subcommittee II of ASTM Committee E10 on Radioisotopes and Radiation Effects.
The specimens were machined from a 96-inch-wide by 72-inch-long by 6-inch-thick plate which was melted using a fine grain practice and a transverse-to-longitudinal rolling ratio of 1:1.
- a. Chemical Analyses (percent)
C Hn P
S Mo Si 0.24 1.34 0.011 0.023 0.51 0.23 4
- b. Heat Treatment The U. S. Steel material was heat treated at the U. S. Steel Homestead District Works as follows:
The 6einch-thick plate was charged into a furnace operating at 1100 F heated at a maximum rate of 63 F per hour, to 1650 F, held at temperature for four hours, and water-quenched to 300 F.
The plate was then recharged into a furnace operating at 700 to 750 F and heated at a maximum rate of 63 F ger hour to 1200 F for six hours.
- The information in this Appendix is from Yanichko, S.
E., "Wisconsin Michigan Power Company Point Beach Unit No.1 Reactor Vessel Radiation Surveillance Program", WCAP-7513, Westinghouse Electric Corporation (June, 1970).
een i
i i
~.
o.
APPENDIX C INSTRLHDRED CHARPY EXAMINATION O
e f
ese e
e
APPENDIX C INSTRUMENTED CHARPY EXAMINATION
SUMMARY
The instrumented Charpy technique was applied to the impact specimen evaluations of the following irradiated pressure vessel materials: two different base metal plates, weld metal, heat-affected zone (HAZ) metal, and ASTM correlation-monitor metal. Because of the limited number of specimens and their need in establishing other parameters such as the upper shelf energy and the NDTT, there were insufficient data available for a complete analysis of the effects of radiation on these materials. However, load-time traces were obtained for all 48 impact specimens. These traces show the change in impact behavior as a function of temperature for all five materials tested.
INTRODUCTION The radiation embrittlement of an operating nuclear pressure vessel is determined by the accelerated irradiation of the original materials as part of a surveillance program. The lifetime of the pressure vessel will depend on the radiation-induced shif t in the ductile-brittle transition temperature as i
measured by the Charpy V-notch inpact test.
Although a few fracture-toughness specimens are also included in some surveillance programs, their number is very limited and the present safety criteria are primarily based on Charpy impact behavior.
The value of irradiated Charpy impact specimens, particularly in present surveillance programs, can be considerably enhanced by the use of the instrumented Charpy test.
The instrumented Charpy test provides a valuable link between the transition-temperature approach and the fracture-mechanics approach to fracture toughness.
A knowledge of the effect of radiation on key metallurgical fracture parameters can be used to accurately predict (1) the radiation-induced shif t in the ductile-brittle ':ransition temperature, and (2) the radiation-inducedchangeinthedynamicfracturetoughness,bd. The results obtained by applying these techniques to the Charpy specimens contained in the first j
surveillance capsule of the Point Beach Unit No.1 reactor are' presented in this section of the report.
C-2
_BACKGR0tMD There are two approaches to determining the effect of radiation on the fracture toughness of pressure-vessel steels: (1) the shif t in the ductile-brittle transition temperature (DBTT)*, and (2) the change in the fracture toughness (either the static fracture toughness K r the dynamic Ic fracture toughness kid). The modern theories of the fracture define key metal-lurgical fracture parameters such as friction stress, grain size, grain-size dependence of the yield stress, and surface energy or plastic work of micro-crack propagation. The effect of radiation on most of these metallurgical fracture parameters has been previously studied, but until recently, the results had not been directly linked with the radiation-induced change in fracture toughness. This recent work established the relationships between the key metallurgical fracture parameters and the DBTI and K,
7 The instrumented Charpy test is an excellent tool for determination of the effects of radiation on the key metallurgical fracture parameters.
This test provides load-time information in addition to the energy absorbed.
The loads involved during impact are obtained by instrumenting the Charpy striker with strain gages so that the striker is essentially a load cell. The details of this technique have been reported previouslyI)
The additional information obtained from the instru=ented Charpy test is the general yield load, P (plastic yielding across the entire cross g
'section of the Charpy spec! men), the maximum load, P,,
the brittle fracture load, P, and the time to brittic fracture (see Figure C-1).
Also, tl.e area p
under the load-time curve corresponds to the total energy absorbed, which is the only data obtained in a normal uninstrumented Charpy test. The instrumented test, however, allows separation of the energy absorbed into (1) the energy required to initiate ductile or brittle fracture (premaximum load energy),
(2) the energy required for ductile tearing (postmaximum load energy), and (3) the energy associated with shear lip formation (postbrittle fracture energy),
l as shown in Figure C-1.
l l
1
- The DBTT should not be confused with the nil ductility transition temperature (NDTT), which is defined by a specific type of test as described in ASm E208.
j
- References at the end of this Appendix.
1 I
C-3 MAXII/.UM LOAD,.Pmo GEt;ERAL YlELD
/ LOAD, PoY N
f yBRITTLE FRACTURE LOAD, Pp POST "MAXIMUM-LOAD"
?
ENERGY 0
\\'
Q:X ENERGY W
POST BRITTLE-FRACTURE PRE"MAXlMUM-LOAD" EllERGY,
1 y
l y y,-
/
i I
1..
Is
.As
,s I:
TitaE TO BRITTLE FRACTURE-Tirne J
e FIGURE C-1.
AN IDEALIZED LOAD-TIME HISTORY FOR A CHARPY IMPACT TEST a
C-4 In a normal Charpy impact study, the energy absorbed is determined as a function of temperature to obtain the Charpy impact curve and the ductile-brittle transition temperature (DBIT). The instru-mented Charpy test also gives the information shown in Figure C-1 as a function of temperature, as shown by the example in Figure C-2, Various investigators ( ~0) have developed theories t. hat permit a detailed analysis of the load-temperature diagram. This diagram can be divided into four regions of fracture behavior, as shown in Figure C-2.
In each region different fracture parameters are involved ( ).
Extended discussions of these fracture parameters can be found in the references indicated above.
In general, the key metallurgical fracture parameters for radiation damage studies are the cicavage fracture stress, o *, and the yield f
strength a. Both of these parameters and the temperature sensitivity of oy can be derived from the results of the instrumented Charpy tests.
The decernination of the cleavage fracture stress o *, requires an evaluation of f
P at the temperature where P is 80 percent of P n e yie M streng6,
i gy p
gy.
o,is calculated from the general yield load, Pgy, and is related to the y
uniaxial tensile strength, 073, by the relation (7)-
~
y3 = 33.3 Pgy.
0 This relation is for a standard Charpy V-notch specimen, is dependent on the flank angle of the notch, and assumes Tresca yield criterion.
g EXPERIMENTAL PROCEDURES The general procedures for the instrumented Charpy test are the same as those for the conventional impact test, and are described in Part A of this report.
The additional data are obtained through a fairly simple electronic configuration, as shown in the schematic diagram of Figure C-3.
9
..,.e em.
4
_,x
C-5 l
Y N N N N N
N N
s%
p:
3 p
me s
F E
G 5
Pr 3
. Pay
.3 2
%E Region i Reglen 2 Region 3 Region 4 Test Tempereture FIGURE C-2 GRAPHICAL ANALYSIS OF CHARPY IMPACT TEST DATA 8
Sep.
--,w-,-
,.,-..,.,-n.,
C-6
(
lI tr "
7-- w I
U Bridge Botence Oscilloscope And 2
o Arrolifier Shunt Triggering Resistance Device L
Hammer r
l i
i l
t FIGURE C-3.
DIAGRAM OF INSTRtl11DITATION ASSOCIATED i
WIDI INSTRUMENTED CHARPY TEST i
i i
4 l
r E
S l
m
,-g-
,,7 7--.
,,-,-,v
t 9
C-7 The striker of the impact machine is modified to make it a dynamic load sensor. The modification consists of a four-arm resistance strain gage bridge positioned on the striker to detect the canpression loading of the striker du' ring the impact loading of the specimen. The compressive elastic strain signal resulting from the striker contacting the specimen is conditioned by a high-gain dynamic amplifier and the output is photographed as it develops on the cathode ray tube of an oscilloscope.
A previously established calibration method ( ) is used to convert the oscillograph into a time-load record.
The time-load history as a function of test temperature forms the basis for further data analysis.
The oscilloscope is triggered by a solid-state device at the correct time to capture the amplifier output signal.
This device consists of a regulated d-c power supply used to operate an incandescent lamp, which acts as a light source. When the Charpy hammer swings by the light source (which is located near the base of the impact tester), the light is reflected from a small nietal mirror located on the side of the hammer. The reflected light is detected by a phototransitor which drives the trigger.
_RESULTS AND DISCUSSION The instrumented Charpy tests were conducted following the procedures
' discussed in Part A of this report.
Specimens were tested from five irradiated materials. These were the pressure-vessel base-metal plates A9811, base metal plate C1423, heat-affected zone metal, weld metal, and AS1H correlation monitor material. The results of these tests and the corresponding load-time records are given in Tables C-1 through C-5.
The tables list the specimen numbers, test temperature, impact energy, general yield load, and maximum load.
It can readily be observed that the features of the load-time or load-deflection traces change as a function of temperature; however, all tests f all into one of the six distinctive notch-bar bending classifications shown in Figure C-4 The pertinent data used in the analysis of each record are the general yield load (Pg), atximum load (P,,x), and f racture load (P ).
The impact energy y
values listed in the tables are those normally obtained frora the impact machine dial.
These values are in exec 11ent agreement with energy values calculated from the area under the load-time curves.
e
C-8 TABLE C-1.
INSDLHENTED CHARPY IMPACT DATA FOR BASE-METAL PLATE A9811 Specimen No.
A2 D,000 Test Temperature, F
-65
~
~
Impact Energy, f t-lb 4.5 General Yield Load, Pgy, Ib fSp00 Maximum Load, P,,x, Ib 3310 0 o 000 2000 Tee, pm i
e i
6 Specimen No.
A5 0,000 Test Temperature, F
-25
-f Impact Energy, f t-lb 5.5 Spoo -
3 General Yield Load, Pgy, Ib i
Maximum Load, Pmax, Ib 3360 e
e i
e i
t 0 0 000 2000 Trneinec i
i i
i i
i i
Specimen No.
A7 O'#
~
~
Test Te=perature, F 5
~
~
Impact Energy, ft-lb 8
[ 5p00 General Yield Load, Pgy, Ib
~
t i
f f
f i
f f
f Maximum Load, P,,, Ib 3310 0
0 000 2000 Tae.psec
.iii
>>iy Specimen No.
A3 10$00 4
Test Temperature, F 26 Impact Energy, f t-lb 22 spo0 General Yield Load, Pgy, Ib 3100 Maximun Load, Pmax, Ib 3540 o
i i
0 000 2000 Twa,put
C-10 FOR BASE-METAL TABLE C-1 (Continued) e i
i i
i i
i i
i i
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i e
i i
A10 0,000 119 B
71 j 5,000 lb 2810 e
i f
3810 0
b, 2000 000 2000 g
K00 Go 5@
T"'i#8 Tee,sses i
i 4
i i
e i
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i A8 D'E
~
~
162
~
a g5,000 87 Ib 3690 g
3690 0
000 0
0 00 200c ' E00 G o ' 5000 Time,sses Tes,suc i
i i
i i
i i
A4 0,000 208
~
33,000 88 f''''-
b 3600 3640 0
000 2M 0
000 2000 Moo 4000 5000 Time, esac Tae,ssac i
i i
i i
i i
i e
i i
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i i
Al O'000
- 1'
~
290
~
~
fh000 90 2460 f
i i.
3480 10 p.,, '
000 0
000 2000 5000 4000
/.00 fee,suc
,._.7.
C-12 TABLE C-2 (Continued)
- i i
e i
i i
e i
i Specimen No.
C7 0,000
~
Test Temperature, F S
Impact Energy, ft-lb 37 5,000 General Yield Load, PGY, ib 3400 Maximum Load, P x, Ib 4NO 0 0 2000 Time, pec i
i i
e i
i e
i i
Specimen No.
C12 m,000 Test Temperature, F 26
~
a Impact Energy, f t-lb 35
} 5,000 General Yield Load, P
, Ib 3320 0
Maximum Load, Pmax, Ib 4200
' y '
' 2000 0
Fee.pm i
i i
i i
i
~
Specimen No.
C11
~
~
Test Temperature, F 33 f 5,000 Impact Energy, ft-lb 38 General Yield Load, Pgy, 1b 3280
~
Maximum Load, P,,x, Ib 4180 0
0 Eco 2000 Tee.sm i
i i
i i
i i
Specimen No.
C8 Tes t Temperature, F 52 m
h' Impact Energy, f t-lb
~
55 Cencral Yield Load, Pgy, ib 3230 1
y '
' gnon Maximum Load, P,,x, Ib 4g 9
C w,,-
C-13 TABLE C-2 (Continued) i i
i i
i i
i i
i Specimen No, C2 0800
~
~
Test Temperature, F 79 Impact Energy, ft-lb 76 5 S00
~
~
Ceneral Yield Load, Pgy, Ib 3110
' y (' 2000
~
g Maximum Load, Pmax, Ib 4200 0
5000 40JO
$000 g
TW. sus i
i i
i i
e i
e i
0,000 Specimen No.
C4 Test Temperature, F 120 a
45 3,000 Impact Energy, ft-lb 101 General Yield Load, PGY, Ib 3030 I.
t g
Maximum Load, Pmax, Ib 4140 0
1000 '
2000 3000 4000 KC0 Twe,suc i
i Specimen No.
C9 0 800 Test Temperature, F 208
~
Impact Energy, f t-lb 119 5p00 General Yield Load, Pgy, Ib 2810 Maximus Load, P,,,, Ib 3940 0
' b'2b'b'b'N 0
Tro,asec i
e e
i e
i i
i 6
0,000 Specimen No.
C6 Test Temperature, F 300 8,
Impact Energy, f t-lb 121 General Yield Load, Pgy, Ib 2560 t
0 Maximum Load, Pmax, Ib 3790 0
000 2000 3000 4000
".000 y g,,,,,,
9
C-14 TABLE C-3.
INSITtt.HENTED CHARPY DiPACT DATA FOR HEAT-AFFECTED ZONE HETAL 1
i i
i i
i 4
6 i
Specimen No.
km4 10,000 Test Temperature, F
-87 e
Impact Energy, f t-lb 5.5 js,000 General Yield Load, Pgy, Ib Maximum Load, P,,x, Ib 3740 0
y '
'2M Time.mc i
i i
6 i
i 10,000 Specimen No, km5 Test Temperature, F
-26 e
f
~
Impact Energy, f t-lb 31 j.
General Yield Load, Pgy, Ib 3600 0
Maximum Load, P Ib 4320 0
000 2M
- max, TW,sm i
i i
e i
i i
e i
Specimen No, hy7 10,000 -
Test Temperature, F 0
~
Impact Energy, f t-lb 33 j $,000
~
General Yield Load, Pg7, Ib 3550
~
Maximum Load, Pau, Ib 4220 0
f f
f f
f f
f 0
000 2000 inaasm i
i i
i i
i i
Specimen No, hms Test Temperature, F 25 8
Impact Energy, f t-lb 20 General Yield Load, PCY, Ib 3520 Maximum Load, Pmax, Ib-3830 0
000 2000
,TN,nec
C-15 TABLE C-3 (Continued) i i
i i
i
'4 i
i i
Specimen No.
W3 10#00 Test Temperature, F 36 j5,000 Impact Energy, ft-lb 99 General Yield Load, Pgy, Ib 3340 Maximum Load, Pmax, Ib 4330 0
0 1000 2@
Tet, me i
a i
i i
i i
e i
Specimen No.
W6 Test Temperature, F 51 m,
~
Impact Energy, f t-lb 83
~
General Yield Load, PGY, Ib 3420 0_
Maximum Load, Pmax, Ib 4470 0
000 2@
7,,,,,
e s
i i
i s
i i
10,000 Specimen No.
W1 Test Temperature, F 79 f
Impact Energy, f t-lb 110 General Yield Load, Pgy, B 3280
,0 000 2@
m 4000 m
Maximum Load, Pmax, Ib 4270 Time.pm i
i i
i
~
Specimen No.
W2
~
Test Terperature, F 208 a'
T
~
('*
~
Impact Energy, f t-lb 105 General Yield Load, pGY, Ib 2840 Maximum Load, P 3790 0
- max, 0
2M M
4M m
Tae,sm l
i C-16 TABLE C-4 INSTRUMENTED CilARPY IMPACT DATA FOR WILD METAL i
i i
i i
i i
i i
Specimen No.
W8 10$00 Test Temperature, F 87
$ $poo Iepact Energy, f t-lb 2.5 General Yield Load, Pgy, Ib Max imum Lo ad, P""*, Ib 2180
.o
'e i
e O
000 2000 Trat,pec i
i i
i e
i e
i i
0,000 Specimen No.
W7 f
Test Temperature, F 31 Impact Energy, ft-lb 6.5 b'"O
~
~
.a i
General Yield Load, Pgy, Ib f
Maximum Load, P,,, Ib 3540 0
0 2000 Time,sut ii.ii i
i e
i 10,000 Specimen No.
W2
,f Test Temperature, F 25 m,
f 15 s,000 Impact Energy, f t-lb 13 General Yield Load, Pgy, Ib 3460
,h,,,,,,,
Maximum Load, Pmax, Ib 3620 c
000 2000 7,..,,,,
iiiii i
i i
i Specimen No.
W4 10p00 Test Tenperature, F 50 Impact Energy, f t-lb 27 p00 General Yield Load, Pgy, Ib 3390 g
Maximum Load, P
, Ib 3920 J
0 0 000 2000 Tr'<,nec
{
l
C-17 TABLE C-4 (C'ontinued) i i
i i
i i
i i
i Speeimen No.
W6 9.000 Test Temperature, F 79 a
Impact Energy, f t-lb 34 5,000 General Yield Load, Pgy, Ib 3320 Maximum Load, Pmax, Ib 3980 0
0 000 2000 Tine, put i
i i
i i
i i
i i
Specimen No.
W5 Test Temperature, F 123 2,
Impac t Energy, f t-lb 46 3
~
~
~
Ceneral Yield Load, Pgy, Ib 3200 t
i t
t t
t t
g Maximum Load, P,,x, Ib 3940 0
000 2000 Tm,pec i
i i
i i
i i
i i
Specimen No.
W1 g,ooo Test Temperature, F 167 A
Impact Energy, ft-lb 54 y 5, con General Yield 1 ad, Pgy, Ib 3100 Maximum Load, Pmax, Ib 3850 t
e t
0 0
gm 4000 5000 Tm,nts i
i i
i e
i i
i Specimen No.
W3 0,000 Test Temperature, F 208 Impact Energy, ft-lb 52
{'5,000 Ceneral Yield Load, PCY, Ib 3050 y
Maximum Load, P
, Ib 3770 0
2m ' MM 0
40M W
i Ten.pnc
(
C-17 TABLE C-4 (C'ontinued) i i
i i
i i
i e
i Specimen No.
W6 10,000 -
Test Temperature, F 79 a
Impact Energy, ft-lb 34 5,000 Cencral Yield Load, Pgy, Ib 3320 Maximum Load, Pmax, Ib 3980 0 0 1000 2000 Tee,pset a
i i
i i
i i
i i
Specimen No.
WS Test Temperature, F 123 8,
Impact Energy, f t-lb 46 3'
~
~
Ceneral Yield Load, Pgy, Ib 3200
~
l I
f f
f 1
1 Maximum Load, P,,x, Ib 3940 0 0 1000 2000 Tee,psee i
e i
i i
i i
i i
Speci=en No.
W1 g,noo.
Test Temperature, F 167 a
Impact Energy, f t-lb 54
$,000 General Yield Load, Pgy, Ib 3100 Maximus Load, Pmax, Ib 3850 a
i i
e i
0 0
1000 2000 3000 4000 5000 Tee,psec i
i i
i i
i i
i i
Specimen No.
W3 10,000 Test Temperature, F 208 Impact Energy, f t-lb 52 j'5,000 General Yield Load, Pgy, Ib 3050 Max inum Load, P "*, Ib 3770 o
0 1000 2000 3000 4000 5000 Tee,nec i
o,
.. - -,. ~
C-18 TABLE C-5.
INSMHENTED CHARPY IMPACT DATA FOR CORRELATION MONITOR MATDtIAL e
i i
i i
i i
i Specimen No.
R6 op00 Test Temperature, F
-25 Impact Energy, f t-lb 3.5
- 8. '3,g General Yield Load, Pgy, Ib Maximun Load, ? ax, lb 2650 m
i e
i e
t t
t i
t-0 0 M
2W Timesw i
i i
i i
i i
e,000 Specimen No.
R1 Test Te=perature, F 50 m,
3 5A00 1epact Energy, ft-lb 10
.5 General Yield Load, Pgy, ib 3420 Maximu:n Load, Pmu, Ib 3480 0
000 2000 Fm,,m i
i i
i i
i i
i Speci=cn No.
R7 ppoo Test Temperature, F 78 Impact Energy, f t-lb 22 e
) 5,000 General Yield Load, Pgy, Ib 3310 Maximum Load, Pma, Ib 3920 0
0 1000 2000 The,ps**
i i
i i
i i
i i
Op%
~
Specimen No.
R5 Test Temperature, F 122
'/
{
{ S.M Impact Energy, f t-lb 20 General Yield Load, Pgy, R 3200 i
t i
t t
i Maximum Load, P,,, Ib 3630 0 0 000 2000 m
4000 x'00 The,mec i
~
C-19 TABLE C-5 (Continued) i e
i i
e i
i i
i Specimen No.
R4 10,000 Test Temperature, F 127 Impact Energy, f t-lb 27
,5,000 General Yield Load, Pgy, ib 3130 j
Maximu:n Load, Pmax, Ib 3790 1
o 0
XXQ 2000 M
M
,0000 Tm,suc i
4 i
i Specimen No.
R2 Test Temperature, F 161 e
Impac t Energy, f t-lb 40 f 5,000 General Yield Load, Pgy, Ib 3100 Maximum Load, P Ib 3920 0
000 m
m 4m Sc00
- max, Tene.put i
i O,000 -
Specimen No.
R3 Test Te=perature, F 188 e,,p Impac t Energy, f t-lb 56 Ceneral Yield Load, Pgy, Ib 3070
\\
0 Maximum Load, P,,x, Ib 4010 0
000 2000 3000 4000 6000 in,, c i
i i
i i
i i
Specimen No.
R8 C:000
~
i Test Temperature, F 236 i
e Impact Energy, it-lb 64 5A00 General Yield Load, Pgy, LB 2880 i.
Maximum Load, Pmax, Ib 3890 0
'N 'm ' 4000 0
000 2000 6000 Tm, rses j
i a
0C 7y LOAD DISPLACEMENT CURVES RAW DATA REMARKS l
Pr Brillie fracture v
i 3
I l
t Deflection i
i Poy Brittle Iroctore i
E l
23 I
I Deflection Poy Brittle freefore fe!! owed by fracture indicative of theor lip formation
}
m Deflection Poy,P,
Stab!e crock prepogotion followed byunstable brillie fracture end frocture indicolive of k
y shtor lip formation h
Defieclion Pg,P,
Stob'e crack propogotion followed by fracture me indicolive of shtar lip formation E
Deflection Poy,P,,
Slobte crock precogerion followed by gross deformotion 10 Dellection FICURE C-4.
THE SIX TYPES OF FRACTURE FOR NOTCllED BAR BDOING
C-21 The Charpy energy curves and the load-temperature information obtained from the instrumented Charpy tests are shown in Figures C-5 through C-9.
These figures illustrate a unique feature of this type of analysis; that is, the determination of a definitive fracture transition temperature by discrimination between fractures occurring below and above general yield (Pg).
This transition is a clear indication of the mechanical properties of the material and does not depend on empirical correlations, as the nil-ductility transition temperature (NDIT) determined by the 30 f t-lb fix temperature does.
It is interesting to note that impact fracture at the 30 f t-lb level corresponds j
to ductile specimen behavior, where considerable work hardening is required to 4
raise the stress at the notch to a value sufficiently above the yield stress for fracture to result.
This is dramatically shown for Specimen W5 in the plot of load versus time in Table C-3.
This specimen was tested at -26 F, which is right at the 30 f t-lb fix (the actual value for Specimen W5 is 31 f t-lb). The plot shows appreciable work hardening af ter yielding rather than a completely brittle failure.
The determination of cleavage fracture stress, o *, required g
evaluation of P at the temperature where P is 80 percent of P Because g
F g.
of the limited number of specimens and the requirements for determination of the
}
NDTT and upper shelf energies, there were insufficient specimens to perform enough tests in the brittle range to well define values of a *.
However, a g
cleavage fracture stress of approximately 270,000 psi was determined for base-metal plate C1423 and the veld material. The cleavage fracture stress was determined for the correlation monitor material as being approximately 260,000 psi.
CONCLUSIONS i
The instrumented Charpy impact test technique was used to study the impact behavior of irradiated pressure-vessel materials and irradiated ASIM I
i correlation-monitor material. Because of the limited number of Charpy specimens, it was not possible to do a complete analysis of the effects of irradiation on these five materials.
However, it was shown that in all four f
materials the nil-ductility transition temperature as determined by the 30 f t-lb
],
fix corresponds to specimen behavior where there is some ductile behavior rather than completely brittle behavior.
i
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c-2a l
6 I
i 4500 4000 O
o
\\
Pmex
.o
\\
o o
B 3500 N
3
\\
N N b 3000 Pg 2500 12 0 10 0 O
-u 80 o
r=
60 g 0
e 8
c 40 "
rr 20 I
t &'
t i
n
-200
-100 0
10 0 200 300 Temperature,F r IGURE C-5.
INSTRtHENTED CliARPY LOAD-TD!PERATURE AND IMPACT D4ERGY-TDiPERATURE CURVES FOR BASE-METAL PLATE A9811
i C-23 l
1 I
i 4500 o
O 4000 Pmax s
\\
\\
o n
h'3500
\\
3 0
3000
- Pay t
2500
' 120 10 0 80 T
O 60 s
?
E 40 "
g 20 r-,
I I
i o
-200
-10 0 0
10 0 200 300 Temperature, F FIGURE C-6 INSTRthtD4TED Cit *RPY LOAD.TDIPERATURE AND IMPACT ENERGY-TDiPERATURE CURVES FOR BASE-METAL PLATE C1423
C-24 l
l i
i i
i l
4500 o
0 0
0 0
4000 b*
s \\
o N
2 e
i 3500 o
.3 i
i Poy i
3000 1
2500 -
l20 D
S 0
100 a T
i 80 E Ew 60 40 o
l 0
20 i
I I
l_
l 0
-200
-10 0 0
iOO 200 300 Temperature, F I
FIGURE C-7.
INSTRUMENTED Cl%RPY LOAD-TDiPERATURE AND IMPACT ENERGY-TDIPERATURE CURVES FOR !! EAT-AFFECTED ZONE HETAL
C-25 t
i i
i i
J 4500 4000 s
P ax m
N N
o N
s t
E 3500 N
.i i
Pay 3000 i
2500 120 100 4
80 s.=
60
~
u
?
8 1
40 "
t c) 20 i
IC-1 I
I 0
-200
-10 0 0
10 0 200 300 Temperature, F FICURE C-8.
INSTRIAIENTED CllAAPY LOAD-TDIPERATURE AND IMPACT ENDLCY-TDIPERATURE CURVES FOR WELD }1ETAL i
C-36 l
I 1
4 l
4500 4000 P
-g mcx NNN 9
0 i3500 N
B P
3000 GY l
2500 120 1
10 0 2
I 4
80 TC
~
60 4 i
o D0 40 j
D a
20 l
l
~
-200
-10 0 0
10 0 2
300 Temperature, F FIGURE C-9 INSTRLHENTED CilARPY LOAD-TDIPERATl'RE AND IMPACT DiERGY-1 TDiPERATURE CURVES FOR ASut CORRELATIO3 HONITOR MATERIAL i
C-27 4
l t
l 1
_ Appendix C References l
(1) Wu11aert, R. A., Ireland, D.R., and Tetelman, A. S., "Radiation Ef fects on the Metallurgical Fracture Parameters and Fracture Toughness i
f of Pressure Vessel Steels", paper presented at the ASTM Annual Meeting, i,
Niagara Falls, Now York, June 29, 1970.
(2) Wu11aert, R. A., "Applications of the Instrumented Charpy Impact Test",
in Innact Testine of }fetals, American Society for Testing and Materials l
Special Technical Publication 466, p 148 (1970).
i (3) Wilshaw, T. R., and Pratt, P. O., "The Effect of Temperature and Strain l
Rate on the Deformation and Fracture of Mild-Steel Charpy Specimens", in Proceedinns of the First International Conference on Fracture, Sendai, Japan, September,1965, 2, p 973.
(4) Tetc1 man, A. S., and McEvily, A. J. R., Fracture of Structural Materials, John Wiley and Sons, Inc., New York (1967).
1 (5) Knott, J.
F., "Some Effects of Hydrostatic Tension on the Fracture Behavior I
of Mild Steel", Ph.D. dissertation, University of Cambridge, Cambridge, Engisnd<(1962),
j (6) Fearnehough, G. D., and Hoy, C.
J., "Mechanism of Deformation and Fracture in the Charpy Test as Revealed by Dynamic Recording of Impact Loads",
Iron Steel Inst.,_202, 912 (1964).
(7) Creen, A. P., and Hundy, B. B., J. Mech. Phys. Solids, 4, 128 (1956).
I I
o (8) Server, W.
L., and Tetelman, A. S., "The Use of Precracked Charpy Specimens 4
i j
to Determine Dynamic Frreture Toughness", UCLA-ENG-7153 (July,1971).
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