Regulatory Guide 1.161
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| Issue date: | 06/30/1995 |
| From: | Office of Nuclear Regulatory Research |
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U.S. NUCLEAR REGULATORY COMMISSION June 1995 REGULATORY GIUIDE
OFFICE OF NUCLEAR REGULATORY RESEARCH
REGULATORY GUIDE 1.161 (Draft was DG-1023)
EVALUATION OF REACTOR PRESSURE VESSELS WITH
CHARPY UPPER-SHELF ENERGY LESS THAN 50 FT-LB.
USNRC nEGUIATOU GUIDES Written comments may be submitted to the Rules Review and Directives Branch, DFRPS, ADM, U.S. Nuclear Regulatory Conrvnsslon, Washing Regulatory Guides are Issued to describe and make available to the ptuIbio ton, DC 20555-0001.
such Information as methods acceptable to the NRO staff for Implement ing specific parts of the Commission's regulations, techniques used by The guides are Issued In the following ten broad divisions:
the staff in evaluating specific problems or postulated accidents, and 1. Power Reactors S. Products data needed by the NRC staff In Its review of applications for permits and 2. Research and Test Reactors 7. Transportation licenses. Regulatory guides are not substitutes for regulations, and com 3. Fuels and Materials Facilities 8. Occupational Health pliance with them Is not required. Methods and solutions different from 4. Environmental and Siting :. Antitrust and Financial Review those set out In the guldes will be acceptable If they provide a basis for the 5. Materials and Plant Protection 10. General findings requIsite to the Issuance or continuance of a permit or license by the Commission. Single copies of regulatory guides may be obtained free of charge by writ ing the Office of Administration, Attention: Distribution and Services Section, U.S. Nuclear Regulatory Commission, Washington, DC
This guide was Issued after consideration of comments received from the 206565-0001; or by fax at (301)415-2260.
public. Comments and suggestions for Improvements in these guldes are Issued guides may also be purchased from the National Technical Infor encouraged at all times, and guides wii be revised,
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CONTENTS
A. INTRODUCTION
.............................................................................. 1 B. DISCU SSION .................................................................................. I
NOMENCLATURE .............................................................................. 2
C. REGULATORY POSITION
....................................................................... 3
1. ACCEPTANCE CRITERIA ................................................................ 3
1.1 Levcl A mnd B Conditions ........................................................... 3
1.2 Leve C Condition ................................................................. 3
1.3 Level D Condition ................................................................. 4
2. ANALYSIS METHODS ................................................................... 4
2.1 Level A mnd B Conditions ........................................................... 4
2.2 Level C Condition ................................................................. 7
2.3 Level D Condition ................................................................. 8
3. MATERIAL PROPERTIES ................................................................ 8
3.1 Welds Made Using Linde 80 Flux .................................................... 9
3.2 Generic Reactor Pressure Vessel Welds ............................................... 9
3.3 Reactor Pressure Vessel Base (Plate) Materials ......................................... 9
4. TRANSIENT SELECTION ............................................................... II
4.1 Plant-Specific Transients .......................................................... I
4.2 Bounding Transients .............................................................. 1 D. IMPLEM ENTATION ........................................................................... 11 REFERENCES ........................................................................................ 12 APPENDIX A: Examples ............................................................................. A-1 APPENDIX B: Computation of Stress Intensity Factors ...................................................... B-l REGULATORY ANALYSIS ............................................................................ R-l m.°
A. INTRODUCTION
provided methods for evaluating the fracture behavior of these materials. Further, Generic Letter 82-26 (Ref. 6) was issued
> Appendix 0. Fracture Toughness Requirements" to to advise licensees of the USI resolutio
n. No new require
10 CFR Part 50, "Domestic .Licensing of Production and ments were implemented as part of the USI resolution.
Utilization Facilities," requires, in part, that the reactor vessel However, neither NUREG-0744 nor Generic Letter 82-26 betline materials ".. must have Charpy upper- shelfnergy contained criteria for demonstrating equivalence of margins of no less than 75 fl-lb (102J) initially and must maintain with Appendix 0 of the ASME Code. Rather, the NRC staff upp*r-shelf energy throughout the life of the vessel of no less asked Section X) of the ASME Boiler Pressure Vessel Code than 50 ft-lb (68J), unless it is demotrated in a manner Committee to develop and suggest to the staff appropriate approved by the Director, Office of Nuclear Reactor Regula criteria.
tion, that lower values of upper-shelf energy will provide In February 1991, the Chairman of the ASME Section margin of safety against fracture equivalent to those required XI Subgroup on Evaluation and Standards provided to the by Appe.ndx 0 of the ASME Code."' Charpy uppcr-shelf NRC staffcriteria that had been developed by members of the energy is defined in ASTM E 185-79 (Ref. 1) and -82 Working Group on Flaw Evaluation (WGFE) and the Working WR2), which are incorporated by reference in Appendix KL Group on Operating Plant Criteria (WGOPC) (Ref 7).
"Reactor Vessel Material Surveillance Program Require Although these criteria did not represent ASME Code criteria, ments to 10 CFR Part 50. This guide describes general they did represent the best opinion of knowledgeable persons procedures acceptable to the NRC staff for demonstrating familiar with the problem and with the ASME Code.
equivalence to the margins of safety in Appendix 0 of the Upon review, the NRC staff found these criteria to be ASME Code (Ref. 3). Several examples using these proc acceptable for denstrating margins of safety equivalent to dures are presented in Appendix A to this guide and in more those in Appendix 0 of the ASME Code (Ref. 3). However.
detail in NUREG/CR-6023 Ref. 4). specific methods for evaluating the criteria still were being This regulatory guide contains information collections developedby the ognizant ASME Code committees. Further, that are subject to the Paperwork Reduction Act of 1980 (44 those efforts were not expected to provide specific guidance U.SC. 3501 et seq.). This regulatory guide has been submit on determining event sequences and transients to be consid ted to the Office of Management and Budget for review and ered, nor were they exected to provide specific guidance on approval of the information collections. These information approprit material properties.
collections and record keeping are needed for demonstrating This guide has been developed to provide comprehen compliance with Appendix 0 to 10 CFR Part 50 for the sive guidance acceptable to the NRC staff for evaluating remaining duratienofthe plants license if Charpy upper-shelf reactorpressure vesses when the Charpy upper-shelf energy energy ofthe materials in the beitline region may drop, or may falls below the 50 ft-lb limit of Appendix G to 10 CFR
have dropped, below the 50 ft-lb regulatory limit Part 50. The analysis methods in the Regulatory Position are The public reporting burden for this collection of based on methods developed for the ASME Code,Section XI,
information is estimated to average 960 hours0.0111 days <br />0.267 hours <br />0.00159 weeks <br />3.6528e-4 months <br /> per response, Appendix K ( 8). The staff has reviewed the analysis including the time for reviewing instructions, semrcing methods in Appendix K and finds that they are echnically existing data sources, gathering and mAinta;nin the data acceptable but are not complete, because Appendix K does needed, and completing and reviewing the collection of not provide information on the selection of transients and ifnnation. Send comments regarding this burden estimate or gives very little detail on the selection of material properties.
any other aspect of this collection of information, including In this regulatory guide, specific guidance is provided on suggestions for further reducing the reporting burden, to the selecting transients for consideration and on appropriate Information and Records Management Branch (T6F33), U.S. material properties to be used in the analyses.
Nulear Regulatory Caomission, Washington DC 20555; and Ductile tearing is the dominant fracture process in the to the Desk Officer, Office of Information and Regulatory upper-shelf region of the Charpy impact energy versus Affairs, NEOB-10202 (3150-0011), Office of Management a .aim fnveor RPV materials. The conditions govern and Budget, Washington, DC 20503. ing cleavage mode-conversion of the ductile tearing process in materials with low Charpy upper-shelf energy are still not
B. DISCUSSION
well understood and are not considered in this regulatory guide.
The problem of evaluating materials that do not satisfy The material property needed to characterize ductile the 50 ft-lb upper-shdf enry requirement was recognized by taring in the analysis methods in this regatory guide is the the NRC staff several years ago and was designated Unre material's J-integral fracture resistance, the J-R curve. This solved Safety Issue A-1 1, "Reactor Vessel Materials Tough curve is a function of the material, the irradiation condition, ness." In 1982, the staffcompleted resolution ofUSI A-l I by the loading rate, and the material temperature. The curve is issuing NUREG-0744, "Resolution of the Task A-1l Reactor detrmined by testing the specific material, uider the condi
"VesselMaterials Toughness Safety Issue (Ref. 5), which tions of interest, in accordance with the American Society for
1.161-1
Testing and Materials Standard Test Method E 1152-87, E' /(_v2)O(ksi).
"Standard Test Method for Determining J-R Curves' (Ref. 9).
Unfortunately, the specific material of interest (Le., the F,. F2, F3 Geometry factors used to calculate the stress material from the beltline region of the reactor vessel under intensity factors (dimensionless).
operation) is seldom available for testing. Thus, testing 2).
programs have used generic materials that are expected to Jww The J-integral from the applied loads (in.-lb/i.
represent the range of actual materials used in fabricating reactor pressure vessels in the United States. Statistical Jm,.ia The material's J-integral fracture resistance (in.
analyses of these generic data have been performed and lb/in.2), J-R curve.
reported in NUREG/CR-5729, -Multivariable Modeling of Pressure Vessel and Piping J-R Data' (Ref. 10). These jai The material's J-integral fracture resistance at a analyses provide a method for determnining the material's J ductile flaw growth of 0.10 in. (in.-lb/'in).
integral fractre resistance that the NRC staff finds acceptable for use in the methods described in this guide. Other methods KI& The mode I stress intensity factor caused by the for determining the material property may be used on an radial thermal gradient through the cladding individual-case basis ifjustified. applied to the vessel inner surface, calculated with no plastic zone correction (ksi Ain.).
NOMENCLATURE
KV The mode I stress intensity factor caused by the The following terms are used in this regulatory guide internal pressure, calculated with no plastic-zone and its equations correction (ksi 4tin.); K,.' and K.m are the axial and circumferential values, respectively.
a The flaw depth, which includes ductile flaw growth (in inches). 4ý K,, calculated with a plastic-zone correction (ksi fin.).
a, The effective flaw depth, which includes ductile flaw growth and a plastic-zone correction (in K1 The mode I stress intensity factor caused by the inches). radial thermal gradient through the vessel wall, calculated with no plastio-zone corroection a*. The effective stable flaw depth, which includes ductile flaw growth and a plastic-zone correction Y.; K. calculated with a plastio-zone correction (ksi (in inches).
'in.).
a,*. The effective stable flaw depth at tensile instabil ity of the remaining ligament, which includes p Internal pressure (ksi).
ductile flaw growth and a plastic-zone correction (in inches). p. The maximum accumulation pressure as defined in the plant-specific Overpressure Protection a. The postulated initial flaw depth (in inches). Report, but not eceding 1.1 times the design pressure (ksi).
2c The total flaw length, which includes ductile flaw growth (in inches). 1* The inner radius of the vessel (in inches).
B, Net-section thickness of the ASTM E 1152-87 SF The safety factor (dimensionless).
(Ref. 9) test specimen used in determining mate rial tearing resistance, J-R curve, behavior (in t The wall thickness of the vessel's base metal (in inches). inches).
Cl ,C2 Coefficients used in the equation for the The sum of the vessel wall thickness, t, and the C3, C4 material tearing resistance, J-R curve. cladding thickness, t.L (in inches).
CR The cooldown rate ('F/hour). tlL The thickness of the stainless steel cladding applied to the vessel inner surface (in inches).
CVN Charpy v-notch upper-shelf energy (fl-lb.).
T Metal tWeraturp, at crack-tip, used in the analy E Young's modulus of elasticity (ksi). sis ('F).
1.161-2
WF The margin factor = 2 standard deviations on test where J,, is the J-integral value calculated for the postu data (dimensionless). lated flaw under pressure and thermal loading where the assumed pressure is 1.15 times the maximum accumulation of A reference material's flow stress, specified as 85 pressure, with thermal loading using the plant-specific heatup ksi in ASME Section XI, Appendix K (Ref.8), on and cooldown conditions. The parameter J,, is the J-integral Charpy upper-shelf energy. characteristic of the material's resistance to ductile tearing Q...), as denoted by a J-R curve test, at a crack extension of
1, The material's yield stress (ksi). 0.1 inch.
v Poisson's ratio (dimensionless), specified as 0.3. 1.1.2 The flaw must be stable under ductile crack growth as given by Equation 2:
C. REGULATORY POSITION
< C'V
(2)
1. ACCEPTANCE CRITERIA 8aa a (with load held constant)
The following criteria are acceptable to the NRC staff at for demonstrating that the margins of safety against ductile fracture are equivalent to those in Appendix 0 to Section III p~fod
=
= lal of the ASME Code. Licensees may follow this regulatory guide to determine the equivalent safety margins, or they may use any other methods, procedures, or selection of materials where J,*. is calculated for the postulated flaw under data and transients to demonstrate compliance with Appendix pressure and thermal loading for all service level A and B
o to 10 CFR Part 50. If licensees choose to follow this conditions where the assumed pressure is 1.25 times the regulatory guide, they must use the acceptance criteria, maximum accumulation pressure, with thermal loading, as analysis methods, material properties, and selection of defined above. The material's J-integral fracture resistance transients as described in this regulatory guide. The accep should represent a conservative estimate of the data for the tance criteria are to be satisfied for each category of transients, vessel material under evaluation (i.e., mean - 2 standard namely, Service Load Levels A and B (normal and upset), deviations). Methods for determining the J-integral fracture Level C (emergenc), and Level D (faulted) conditions. These resistance, J-R curve, are discussed in Regulatory Position 3 service load levels are described in Standard Review Plan of this guide. Methods for determining the appropriate service
3.9.3 (Ref 11). Because of differences in acceptable outcome level conditions are discussed in Regulatory Position 4 of this during the various sevice load levels, different criteria have guide.
been developed for Levels A and B, C, and D.
1.2 Level C Condition
1.1 Level A and B Conditions When the Charpy upper-shelf energy of the base metal When the upper-shelf Charpy energy of the base metal is less than 50 ft-lb, postulate both axial and circumferential is less than 50 ft-lb, postulate both axial and circumferential interior flaws and use the toughness properties for the corre interior flaws and use the toughness properties for the corre spending orientation. When the Charpy upper-shelf energy of sponding orientation. For a weld with Charpy upper-shelf any weld material is less than 50 ft-lb, postulate an interior energy less than 50 fl-lb, postulate an interior surface flaw surface flaw with its major axis oriented along the weld of oriented along the weld of concern and orient the flaw plane concern and the flaw plane oriented in the radial direction.
in the radial direction. Postulate a semi-elliptical surface flaw Consider postulated surface flaws with depths up to one-tenth with an aht = 0.25 and with an aspect ratio of 6-to-I surface the base metal wall thickness, plus the clad thickness, but with length to flaw depth. A smaller flaw size may be used on an te total depth not to exceed 1.0 inch (2.54 cm) and with an individual-case basis ifjustified. Two criteria must be satisfied aspect ratio of 6-to- I surface length to flaw depth. A smaller as described below. The maximum accumulation pressure, maximum flaw depth may be used on an individual-cas basis discussed below, is the maximum pressure defined in the Over if justified. For these evaluations, two criteria must be Pressure Protection Report that satisfies the requirement of satisfied.
Section III, NB-731 1(b), of the ASME Code (Re 12).
11.1 The crack driving force must be shown to be less
1.1.1 The crack driving force must be shown to be less than the material toughness as given by Equation 3:
than the material toughness as given by Equation 1:
(1) aWpped < Jo.1 (3)
J~vpied < J0.1
1.161-3
where J., is the J-integral value calculated for the postu including stable tring, should not be greater than 75% of the lWted flaw in the beltline region of the reactor vessel under the vessel wall thickness, and the remaining ligament should be governing Service Level C condition, with a safety factor of safe from tensile instability. The material's J-integral fracture
1.0 on the applied loading. J01 is the J-integral characteristic resistance should reflect a best estimate, Le., the mean value, of the material resistance to ductile tearing (JQ..), as denoted of the data representative of the vessel material under K
by a J-R curve test, at a crack extension of 0.l inch. evaluation.
The J-integral resistance versus crack growth, J-R curve,
1.,2 The flaw must also be stable under ductile crack is discussed in Regulatory Position 3 of this guide. Methods growth as given by Equation 4: for determining the appropriate service level conditions are discussed in Regulatory Position 4 of this guide.
(4)
2. ANALYSIS METHODS
8a aa (with load held constant) The analysis methods described in this guide are at acceptable to the NRC staff for evaluating the criteria de scribed above. Other methods may be used ifjustifled on a i'amp~~ = Jtmlt,1 case-by-case basis.
where Jo is calculated for the postulated flaw under the 2.1 Level A and B Conditions governing Service Level C condition, with a safety factor of
1.0 on the applied loading. The material's J-integral fracture The acceptance criteria discussed in Regulatory Position resistance hould represent a conservative estimate of the data 1.1 for Level A and B conditions involve a comparison of the for the vessel material under evaluation (i.e., mean - 2 applied J-integral to the material's J-integral fracture resis standard deviations). The J-integral resistance versus crack tanc at a ductile flaw extension of 0.1 inch and a determina growth, J-R curve, is defined in Regulatory Position 3 of this tion that this flaw would be stable under the applied loading.
guide. Determination of the appropriate service level condi Procedures are detailed below for (1) calculating the applied tions is discussed in Regulatory Position 4 of this guide.
J-integral for Service Levels A and B flaws and loading conditions and (2) determining that the slope of the material's
1.3 Level D Condition J-integral resistance curve is greater than the slope of the applied J-integral versus crack depth curve at the equillritum When the Charpy upper-shelf energy of the base metal point on the J-R curve where the two curves intersect, as is less than 50 ft-lb, postulate both axial and circumferential illustrated in Figure 1.
interior flaws and use the toughness properties for the corre spending orientation. When the Charpy upper-shelf energy of 2.1.1 Calculation of the Applied J-Integral any weld material is less than 50 ft-lb, postulate an interior The calculation of the applied J-integral consists of two semi-elliptic surface flaw with the major axis oriented along steps: Step 1 is to calculate the effective flaw depth, which the weld of concern and the flaw plane oriented in the radial includes a plastic-zone correction, and Step 2 is to calculate direction. Consider postulated surface flaws with depths up to the J-integral for small-scale yielding based on this effective one-tenth the base metal wall thickness, plus the clad thick flaw depth.
ness, but with total depth not to exceed 1.0 inch (2.54 cm) and with an aspect ratio of 6-to-I surface length to flaw depth. A
smaller maximum flaw depth may be used on an individual Step I
case basis if justified. For an axial flaw with depth'a equal to (0.25t + 0. 1 in.),
For these evaluations, the postulated flaw must be stable calculate the stress intensity factor from internal pressure, p, under ductile crack growth as given by Equation 5: with a safety factor, SF, on pressure equal to 1.15, using Equation 6:
(5)
K F=(s)
p, [I + (Rl11] (7c)°3 F, (6)
(with load held constant)
at F, = 0.982 + 1.006(a/t) 2 Jw1'ppWia = J.mWAial where J.3P is calculated for the postulated flaw under the This equation for K,"" is applicable to 0.05 - a/t < 0.50, and governing Service Level D condition, with a safety factor of it includes the effect of pressure acting on the flaw faces.
1.0 on the applied loading. Additionally, the flaw depth,
1.161-4
E
0)
4)
S -- Evaluation Point Crack Extension, Aa Figure 1. Comparison of the Slope of the Applied J-Integral and J-R Curve.
1.161-5
For a circumferential flaw with depth 'a equal to (0.25t limiting cooldown rate, including the contributions of cladding
+ 0. 1 in.), calculate the stress intensity factor from internal to thermal stress and the thermal stress intensity factor. For pressure, p., with a safety factor, SF, on pressure equal to this alternative analysis method (also described in Reference,
1.15, using Equation 7: 4), the main features for computing KY,and Kk., which are applied in examples in Appendix A, are given in.
K4'- =($F)p. [ I +(R,/(2t1))] ( a)°0'SF, 7 Appendix B.1- The limiting condition should be determined for the transient time at which the material's J-R curve will be greater than or equal to the Jbw for evaluating Equations 1 and 2. The main steps are:
F2 = 0.885 + 0.233 (aft) + 0.345(a/t)62 a. Determine the temperature gradient across the vessel This equation for Kip' is applicable to 0.05 :c a/t s 0.50, wall thickness, in 10 to 20 time steps over thefull and it includes the effect of pressure acting on the flaw faces. duration of the tra*sLient; and compute the corresponding For an axial or circumferential flaw with depth 'a' equal thermal stress histozy, taking into account the cladding to (0.25t + 0.1 in.), the "steady-state" (time independent) thickness, t,.
sess intensity factor from radial thermal gradients is obtained b. For each time step. compute K. and uL values as a by using Equation 8: function of the crack depth in the range 0.05 s a/t s
0.5.
c. For Equation 1. calculate the pressure-induced K, and KIt = ((CR)10oo)t" F3 (8) the J... using Equations 9 and 10, at a crack-tip depth of(0.25t' + 0.1 in.) for each time step.
F3 = 0.69 + 3.127(a/t) - 7.435(a/62 + 3.532(a/t)3 d. Use Step a to find crack-tip temperature history at each time step. See Figure A-I in Appendix A for an This equation for Kit is valid for 0.2 :g a/t 0.50, and 0 s CR
example.
g 100F&r. This equation does not include the contribution to e, For a given material condition, determine the J-R values K, from the cladding thickness, t ... If the steady-state values at the crack extension of 0.1 inch by using the crack-tip of thermally induced K,, are used, the material J-R curve temperature history from Step d. See Figure A-2 in should correspond to the temperature at the beginning of the Appendix A for an example.
transient, when a uniformly high temperature is present across f Compare the material's J-R values as a function of time the vessel wall thickness, leading to the lowest J-R curve. The above Kjt expression can be replaced with an improved in Step e with the Jý, values in Step c. See Figure A-2 K
in Appendix A for an example. The time at which the accuracy solution if an appropriate justification is provided. J-R value is just equal to the J,,w determines the Calculate the effective flaw depth for small-scale critical condition for evaluating Equation 1.
yielding, a,, using Equation 9: g. At the time determined in Step Z evaluate Equation 2 to verify the stability of the predicted flaw growth.
a, = a + (T) [( t*_..
7[ U (9) 2.1.2 Evaluation of Flaw Stability Flaw stability is evaluated by a direct application of the Step 2 flaw stability criterion given by Equation
2. The applied J
For an axial flaw, calculate the stress intensity factor integral is calculated for a series of flaw depths corresponding from internal pressure for small-scale yielding, C, by to increasing amounts of ductile flaw growth. The applied substituting a. in place of 'a' in Equation 6, including the pressure, p, is set equal to the maximum accumulated pressure equation for F,. For a circumferential flaw, calculate K; by for Savice, Level A and B conditions, ps, with a safety factor, substituting a. in place of 'a' in Equation 7, including the SF, equal to 1.25. The applied J-integral for Service Level A
equation for F2. For an axial or circumferential flaw, calculate and B conditions may be calculated using Equations 6 through the sress intensity factor from the radial thermal gradients for 10. Each pair of the applied J-integral and flaw depth is small-scale yielding Ký, by substituting a, in place of 'a' in plotted on a crack driving force diagram to produce the Equation 8, including the equation for F3.
The J-integral from the applied loads for small-scale yielding is given by Equation 10: 'The equations provided in Appendix B may be used if the transient temperature hstoy can be approxmatedI adequaty by either an exponential or a polynomial equatio. IVit cannot be approximated J,,,,~a = 1000(K,,a.4 IE' (10) adequately, a mom rigorous approach should be used.
Alternatively, in place of the steady-state Equation 8, a Te omer code egmviuAppeandixB is for general illustration licensees thermal transient stress analysis may be performed for the assume responslity for the correctness of the computer codes ty use.
1.161-6
applied J-integral curve as illustrated in Figure 1. The mate riars J-R curve also is plotted on the crack driving force K4p -=(S*)p,(. I +R/(2t)](a)O.SF 2 (12)
diagram. Flaw stbiity at a given applied load is demonstrated if the slope of the applied J-integral curve is less than the slope of the material's J-R curve at the equilibrium point on F2 = 0.885.0233(a/ft)+0.345(a/t)2 the J-R curve where the two curves intersect.
2.2 Level C Condition These equations for Kip'"w are valid for 0.05 s at' s 0.5, and include the effect of pressure acting on the flaw faces.
The acceptance criteria discussed in Regulatory Position If it can be demonstrated that the actual cooldown rate
1 for Service Level C conditions are similar to those for could be bounded by a "constant" cooldown rate, for each Service Levels A and B, with the exceptions of the crack size crack depth the stress intensity factor arising from radial to be considered and the safety factor applied to the pressure thermal gradient, including cladding effects (see Example 4 in loading. For Service Level C conditions, flaw sizes up to Appendix A) is given by Equation 13:
one-tenth the base metal wall thickness, plus the clad thick ness t.L but with a total depth not to exceed 1.0 inch (2.54 cm), are to be considered. A safety factor of 1.0 is used for K,-[-0.012771 *0.549525(- R)-0.611352( )2
1000 1000
both pressure and thermal loading. As with the Service Level +(0.565199,0.046752(.-2-))( 1-.95371("y A and B criteria, for Service Level C it must be demonstrated 1000 t I (13)
that the applied J is less than the material's fracture resistance *1.6287(-a1(t*P
at a crack extension of 0.1 inch, and that the flaw must be t stable under the applied loading.
Procedures are described below for (1) determining the This equation is applicable to 0.05 < a&t' r. 0.5, and 100 g CR
applied J-integral for Service Level C flaw and loading < 600TFhour. The CR values less than 100"F/hour are conditions and (2) determining that the slope of the material's covered under Service Levels A and B (see Equation 8). The J-integral fracture resistance, J-R curve, is greater than the cladding thickness ist. -51l6 in., R, = 86.875 in., base metal slope of the applied J-integral versus crack depth curve. thickness t = 8.625 in., and RA' ratio = 9.72. Details of the analysis results are given in Appendix A. Equation 13 is based
2.2.1 Calculation of the Applied J-Integral on the current state of knowledge on K solutions for 6:1 The calculation of the applied J-integral consists of two asect-ratio flaws subjected to non-uniform stress gradients in steps: Step I is to calculate the effective flaw depth, which the crack-depth direction. The above I. expression can be replaced with an improved accuracy solution if an appropriate includes a plastic-zone correction, and Step 2 is to calculate the J-integral for small-scale yielding based on this effective justification is provided.
flaw depth. Calculate the effective flaw depth for small-scale yielding, a,, using Equation 14:
Step )
Postulate a series of flaws with depths ranging up to cladding thickness plus 0. 1 times the base metal wall thick = a + (-L) I(K Ia Kft) 12 (14)
ness, but not exceeding 1.0 inch (2.54 cm). The number of flaws and the specific flaw sizes to be postulated should be Step 2 m icrient to determine the peak value of the applied J-integral For each flaw size considered, calculate the stress over this size range. For each of these postulated flaws, the intensity factor arising from internal pressure for small-scale analysis flaw size 'a' should be the sum of the postulated flaw yielding, IC by substituting a. in place of 'a' in Equation 11 size plus 0. 1-inch ductile crack extension. For axial flaws, at for the axial flaws and in Equation 12 for the circumferential each analysis flaw size, calculate the stress intensity factor flaws. Similarly, calculate the stress intensity factor arising arising from internal pressure, p., with a safety factor, SF, on from radial thermal gradients for small-scale yielding, Y, by internal pressure equal to 1.0, using Equation 11: substituting a. in place of 'a' in Equation 13. The J-integral arising from the applied loads for small-scale yielding is given K4',"'=(sF9p. [I÷+(R/t) ] (7raf5F, 11 by Equation 15:
F1 =0.982+1.006(alt'?;with 0.05<a/tr'0.5 JWpphtd = 1000(K: 7 K)2 P+ IE / (15)
For circumferential flaws, at each analysis flaw size In an actual transient the cooldown rate initially may calculate the stress intensity factor arising from internal vary sigoificantly with time. Therefore, transient-specific peak pressure, p. with a safety factor, SF, on pressure equal to 1.0, thermal stress-induced KCH and K,4& computations may be
"usingEquation 12: necessary. If so, in place of Equation 13, a thermal transient
1.161-7
stress analysis may be performed for the specific transient, 2.3 Level D Condition including the contrinbutions of cladding to thermal stress and the stress intensity factor. For this alternative analysis method The acceptance criteria discussed in Regulatory Position the main features for computing K. and I., which are applied on examples in Appendix A, are given in Appendix I for Level D Service Conditions involve only the stability of K
the postulated flaws. Additionally, the stable flaw depth must B.u The limiting condition should be determined for the not exceed 75% of the vessel wall thickness, and the remain transient time at which the materiars resistance (I-R curve) ing ligament must be safe from the tensile instability.
will be greater than or equal to the J,, for evaluating Stability of ductile crack extension is demonstrated for Equations I and 2. The main steps are: Service Level D in the same manner used for Service Level C.
However, the material properties should represent only the a. Determine the temperature gradient across the vessel best estimate (i.e, mean value) of the J-R curve for the vessel wall thickness, in 10 to 20 time steps over the full material under evaluation.
duration of the transiet, and compute the corresponding Tensile stability of the remaining ligament is conserva thermal stress history, taking into account the cladding tively demonstrated if Equation 16 is satisfied.
thickness, tI.
b. For each time step, compute Ka and KkL values as a Ofa (16)
function of the crack depth in the range 0.05 : a/t' <
0.5.
c. For Equation 1,calculate the pressure-induced Ky. and Where, from Reference 13, for a semi-elliptical flaw, the J.P.., using Equations 14 and 15, at a crack-tip depth of ((0.( t + tý, + 0.1 in.) < 1 in.) for each time a** - [a*(l - (1 + 2c'A 2)4 )] / (1 - (a*/A){1 + 2c"A2)-]
step.
d. Use Step a to find crack-tip temperature history at each time step. See Figure A-i in Appendix A for an
3. MATERIAL PROPERTIES
example.
e. For a given imterial condition, dtermine the J-R values The statistical analyses reported in Reference 10
at the crack extension of 0.1 inch by using the crack-tip addressed a broad range of materials and conditions For the temperature history from Step d. See Figure A-2 in purposes of this guide, the NRC staff has concluded that only Appendix A for an example. the ASTM E 1152-87 (Ref 9) definition of the J-integral f. Compare the material's J-R values as a function of time fracture resistance curve should be used. This determination \1 in Step e with the J3 values in Step c. See Figure A-2 requires that a test specimen's net thickness, B%, be specified.
in Appendix A for anexample. The time at which the Smaller specimens typically produce more conservative J-R value is just equal to the J, determines the (lower) J-R curves than larger specimens. However, larger critical condition for evaluating Equation 1. specimens are needed to provide large amounts of crack g. At the timedetennined in Step f, evaluate Equation 2 to growth needed in evaluating certain stability criteria described verify the stability of predicted flaw growth. in Regulatory Position 2 of this regulatory guid
e. The NRC
staff recommends the test specimen's net-section thickness, B,
2.2.2 Evaluation of Flaw Stability to be 1.0 inches (2.54 cm) for determining the J-integral Flaw stability is evaluated by a direct application of the resistance curve using the methods specified in Regulatory flaw stability criterion given by Equation 4. The applied J Position 3. This is a reasonable compromise and slightly integral is calculated for a series of flaw depths corresponding simplifies the equations for the material J-ft curve. The to increasing amounts of ductile flaw growth. The applied neutron fluence attenuation at any depth in the vessel wall pressure, p, is set equal to the peak pressure for the Service (such as near the crack tip) should be determined using Level C transient under consideration with a safety factor, SF, Regulatory Guide 1.99 (Ref 14).
equal to 1.0. The applied J-integral for Service Level C This guide provides methods for determining the I
conditions may be calculated using Equations I I through 15. integral fracture resistance of three classes of materials: welds Each pair of the applied J-integral and flaw depth is plotted on Sm factured with Linde 80 welding flux, generic welds used a crack driving force diagram to produce the applied J-integral in fabricating reactor pressure vessels, and plate materials curve as illustrated in Figure 1. The materiars J-R curve also (low and high toughness). The J-R curves for plant-specific is plotted on the crack driving force diagram and intersects the materials may be used if justified on a case-by-case basi.
abscissa at the initial flaw depth, a,. Flaw stability at a given Otherwise, the material's J-integral fracture resistance may be applied load is demonstrated if the slope of the applied J determined from Equation 17, developed in Reference 10:
integral curve is less than the slope of the material's J-R curve at the equilibrium point on the J-R curve where the two curves intersect. J4=(0) {C)(Aa) 0 exp[C3 (Aap])C 4
(17)
1.161-8
The coefficients in Equation 17 for each material type are sulphur content seeins to be areasonable indicator of the plate discussed below. As noted earlier, the net-section thickness, tougnss, with a "higher" sulphur content indicating "lower'
B., of ASTM E 1152-87 (Ref. 9) compact-tension (CT) fracture toughness ( 17). A sulphur content ofO.018 wt-%
specimem to be considered is specified as I inch. In addition is a good demarcation for high- and low-toughness values.
to the Charpy (CVN) models discussed in this guide, Refer Because of the low-toughness plate issue, and because ence 10 contains two other models, namely the Copper ofthe relat*ivy sparse data base that could be used to estimate Fluence (Cu-ft) models and the pre-inradiation Charpy the firature toughness forthese materials, a fracture toughness (CVN) models, which may be used to determine the mate model is only provided for high-toughness plate materials. If rial's J-R curves. the sulphur content of the plate is less than 0.018 wt-%, the plate models described in Reference 10 may be used. How
3.1 Welds Made Using Linde 80 Flux ever, if the sulphur content is greater than or equal to 0.018 wt-%, justification sbxzld be provided for use of the models in For analyses addressing Service Levels A, B, and C, a Reference 10. Factors that might justify use of these high conservative representation of the J-R curve is obtained by toughness models could include information about the year of setting the margin factor, MF = 0.648. For analyses addressing manufacture of the plate and any special thermo-mechanical Service Level D. set MF = 1.0. processing that would serve to improve the fracture toughness of the plate. If adequate justification cannot be provided, a Cl = exp[-3.67 +1.45 ln(CVh) -0.00308T7 low-toughness plate model should be developed and used.
(18) The CVN value should be for the proper orientation of the plate material (see Figure 2). For example, for axial flaws C2 = 0.077 + 0.116 InCl (19) the CVN value for the L-T (strong) orientation in the vessel wall should be used. Similarly, for circunferential flaws the C0 = -0.0812 - 0.0092 lnCI (20) CVN value for the T-L (weak) orientation should be used. In many cases, the CVN values for both orientations may not be C4 = -0.5 (21) known. ff the CVN value for the T-L (weak) orientation is not available, the L-T (strong) orientation CVN value may be
3.2 Generic Reactor Pressure Vessel Welds multiplied by a factor of 0.65 (Ref. 18) to obtain the CVN
value for the T-L (weak) orientation. However, if the CVN
For analyses addressing Service Levels A, B, and C, a value for the T-L (weak) orientation is known and the L-T
conservative representation of the J-R curve is obtained by (sdrug) orientation is to be estimated, the CVN value for the setting the margin faictor, M - 0.629. For analyses addressing L-T (strong) orientation is assumed to be the same as that of Service Level D, set M - 1.0. the T-L (weak) orientation.
C1 = exp[-4.12+1.49 ln(CYh)-0.00249T (22) 3.3.1 High-Toughness Model (S < 0.018 Wt-%)
For plate material with sulphur content greater than C2 = 0.077 + 0.116 InCl (23) 0.018 wt-%, the use of this model should be justified as discussed above.
C3 = -0.0812 - 0.0092 InCI (24) For analyses addressing Service Levels A, B, and C, a conservative representation of the J-R curve is obtained by C4 = -0.5 (25) setting the margh fctor, MW = 0.749. For analyses addressing Service LevelD, set MF - 1.0.
3.3 Reactor Pressure Vessel Base (Plate) Materials The elastic-plastic fracture toughness of plate materials may be relatively high or quite low, depnding on a variety of CI = exp[-2.44+1.13 ln(C/WA)-0.0027771 (26)
chemical, metallurgical, and thermo-mechanical processing variables. The statistical analyses reported in Reference 10 C2 = 0.077 + 0.116 InCl (27)
included only materials that exhibited a J-R curve with a significantly rising slope, ic., the higher toughness materials. C0 = -0.0812 - 0.0092 InCl (28)
However, test results reported in NUREG/CR-5265, "Size Effects on J-R Curves for A-302B Plate" (Ref. 15), clearly show J-R curves with very little, if any, increase in slope. C4 = -0.409 (29)
References 15,16, and 17 provide some insight into the nature ofthe low toughness issue for the plate materials. While there are several variables that influence the fratr toughness,
1.161-9
DEFINITION OF ASME AND ASTM ORIENTATIONS
"WEAK" DIRECTION "STRONG" DIRECTION
ASME TRANSVERSE ASME LONGITUDINAL
K
Figure 2. DefInition of the ASNM and ASIh Flaw Oricntations in an RPV.
1.161-10
3.3.2 Low-Toughness Plate (S k 0.018 Wt-%) ATWS in currently operating light-water-reactor (LWR)
For analyses addressing materials with a sulphur content vessels in the United States is not found to be a dominant greatcr than 0.018 wt-%, the J-R curve data are scarce. Very transient with respect to the low Charpy upper-shelf energy
> limited J-R data for a 6-inch-thick specimen (ASTM 6T CT issue, and no further action is necessary with respect to at 180I' temperature) from an A-302B plate in the T-L ATWS. However, for designs other than the currently operat (weak) orientation, avaiable in NUREOGCR-5265 (Ref 15), ing LWR vessels in the United States, ATWS could become may be used with adjustments for the specimen temperature a dominating transient, and as such needs to be considered as and CVN value (Ref. 19), or a material-specificjustification a Service Level C transient for further evaluation. A plant should be provided to support the use of other data. For specific justification should be provided for consideration of analyses addressing Service Levels A, B, and C, a lower such designs at another service load level. For such designs, bound reptwentation (mean - 2 standard deviations) of the J-R lioensees should consider the assumptions used in the generic curve should be used. For analyses addressing Service Level analyses of Reference 4 to be sure that they are bounding for D, the mean value of the J-R curve should be usecd theirplant-specific applications. If these generic analyses are Additional J-R curve test data for the low-toughness not bounding, plant-specific analyses should be performed.
A302B plate material are presently being generated. Regula tory guidance will be updated, ifjustified, based on the results 4.1 Plant-Specific Transients obtained from the test data collected for J-R curve in low To provide reasonable assurance that the limiting toughness plate material. service loading conditions have been identified on a plant specific basis, the Service Level C and D design transients and
4. TRANSIENT SELECTION events that are necessary to demonstrate compliance with Standard Review Plan 3.9.3 (Ref. 11) should be used.
Selection of the limiting transients for Service Levels C When Ihis transient list isnot available or isincomplete, and D is a key aspect of evaluating the integrity of reactor the most complete list oftransicnts for these service levels that pressure vessels that contain materials with Charpy upper is available for similar plant designs should be used. Typi shelf energy less than 50 fl-lb. Generally, Service Levels A cally, the most complet list of transients would be for the and B are limiting. However, there may be plant-specific later-vintage plants from a particular vendor. This list should consicerations that make Service Levels C or D controlling for be reviewed, and the limiting transients for the reactor vessel ductile fracture. being analyzed should be defined. Once the transients are To provide reasonable assurance that the limiting defined, system-levl thermal-hydraulic analyses should be service loading conditions have been identified, either oftwo performed to determine the limiting presmre-temperature approaches may be used: a plant-specific transient evaluation time history for each transient being considered. This history or a generic bounding analysis. It should be noted that plants provides the input to the analyses described in this guide.
may be grouped and limiting transients for these groups may be determined. The plant-specific transient evaluation is the 4.2 Bounding Transients preferred approach. However, since some licensees may not When the plant-specific transients are not available or have the specific transient infoafnaion needed for this analysis, when developing or updating the pressure-temperature-time a conservative "bounding" anasis may be performed for each histoy would be an undue burden, a conservative "bounding"
service level. Specific guidance for each of these approaches pressure-tenmprature-time history may be used. This history is provided below. shxmld anticipate a pressure equal to the shut-off head for the As described in the Discussion section of this guide, high-pressure injection system and a cooldown rate of 400OF
ductile tearing is the dominant fraure process in the upper per hour for Service Level C and 6001' per hour for Service shelf region, and the possibility of mode-conversion to Level D. These values are based on the NRC staffs experi cleavage (brittle) fracture is not considered in this regulatory ence in performing the bounding analyses (for examples, see guide. The analyses using these bounding transients need only Appendix A of this regulatory guide and Reference 4).
address the transient from its beginning to the time at which Altematives to these cooldown rates may be used ifjustified the metal at the tip of the flaw being analyzed reaches a by the plant-specific safety-injection flows and temperatures.
tm rau equivalent to the adjusted RT. plus 500F. In this regulatory guide, an adusted RTmr plus 50WF (which typically
D. IMPLEMENTATION
represents the low-temperatu overpressure protection TIhe purpose of this section is to provide information to systen's enabling temperature) is taken as the lower tempera applicants and licensees regarding the NRC staffs plans for ture limit for upper-shelfbehavior. using this regulatory guide.
This regulatory guide states that licensees should Except in those cases in which an applicant or licensee consider a spectrum of transients, including ATWS (antici proposes an acceptable alternative method for complying with pated transient without scram). Although ATWS is not a specified portions of the Commission's regulations, the design basis transient, for compliance with Appendix G to 10 methods descnled in this guide reflecting public comments CFR Part 50 it was considered in Reference 4 for evaluation will be used by tlhe NRC staff in the evaluation of applications of low upper-shelf energy materials. Based on the generic fornew licenses and for evaluating compliance with Appendix analyses in Refermne 4 and additional staff calculations, Gto 10 CFR Part 50.
1.161-11
REFERENCES
1. American Society for Testing and Materials, "Standard Practice for Conducting Surveillance Tests for Light-Water Cooled Nuclear Power Reactor Vesses," ASTM E 185-79, July 1979.'
2. American Society for Testing and Materials, "Standard Practice for Conducting Surveillance Tests for Light-Water Cooled Nuclear Power Reactors, ASTM E 185-82, July 1982.'
3. American Society ofMechanical Engineers,Section XI, Division 1, "Rules for Inservice Inspection of Nuclear Power Plant Components,* of the ASVE BoilerandPressureVessel Code, New York, through 1988 Addenda and 1989 Edition. 2
4. TL Dickson, Genric Analyses for Evalualion ofLow Charpy Upper-Shelf Energy Effects on Safety Margins Against Fracture of Reactor Pressure Vessel Materials," USNRC, NUREG/CR-6023, July 1993.3
5. R. Jdmson, "Resolution ofthe Task A-II Reactor Vessel Materials Toughness Safety Issue," USNRC, NUREG-0744, Volume I (Revision 1) and Volume 2 (Revision 1), October 1982.
6. Generic Letter No. 82-26, "NUREO-0744 Rev. 1; Pressure Vessel Material Fracture Toughness," Issued by Darrel G. Eisenhut, Director, Division of Licensing NRR, USNRC, November 12, 1982.!
7. Letter from Warren H Bamford, Chairman of the ASME Subgroup on Evaluation Standards for ASME Section XI,
to James E. Richardson, USNRC, Subject: Response to NRC Request, A-I l Issue, February 20, 1991.!
8. American Society of Mechanical Engineers, Assessment of Reactor Vessels with Low Upper Shelf Charpy Impact EncrU Levels," Appendix K, A93, pp. 482.1-482.15,Section XI, "Rules for Inservice Inspection of Nuclear Power Plant Components," 1992 Edition, 1993 Addenda, New York, December 1993.
9. American Society for Testing and Materials, "Standard Test Method for Determining J-R Curves," ASTM E 1152-87, May 1987.1
10. ED. Eason, J.R Wright and BE. Nelson, 'Multivariable Modeling of Pressure Vessel and Piping J-R Data," USNRC, K,
NUREG/CR-5729, May 1991.3
11. AW. Srkiz, "ASME Code Class 1. 2, and 3 Components, Components Supports, and Core Structures," Revision I
to Appendix A to Section 3.9.3 of NUREG-0800, "Standard Review Plan for the Review of Safety Analysis Reports for Nuclear Power Plants," pages 3.9.3-12 to 3.9.3-20, April 1984.'
12. American Society ofMechanical Engineers,Section III, "Nuclear Power Plant Components," of the ASME Boilerand PressureVessel Code, New York, through 1988 Addenda and 1989 Edition.?
13. J*M. Bloom, "Validation of the Deformation Plasticity Failure Assessment Diagram (DPFAD) Approach - The Case of An Axial Flaw in a Pressurized Cylinder," Transactions of the ASME, JournalofPressure Vessel Technology, Volume 112, pp. 213-217, 1990.
2, May 1988.4
14. USNRC, "Radiation Embrittlement of Reactor Vessel Materials," Regulatory Guide 1.99, Revision
15. A.L. Hiser and J.B. Terrell, "Size Effects on J-R Curves for A-302B Plate," USNRC, NUREG/CR-5265, January
1989.
'Copies may be obtained from the American Society for Testing and Materials, 1916 Race Street, Philadelphia, PA 19103.
'Copies may be obtained fiam the American Society of Mechanical Enginer, 345 East 47th Street, New York. NY 10017.
'Copies am available forinata or copyingfra fee fion the NRC Publio Document Room at 2120 L Street NW., Washingtm, DC; Ie PDR's mailing a&de is Mail Stop LLU Washington DC 20555; telephone (202) 634-3273, fax (202) 634-3343. Copies may be purchased at cunrent rates firom the U.S. Government Printing Office, Post Office Box 37082, Washington, DC 20013-7022 (telephone (202) 512-1S00) or from the National Technical Information Service by writing NTIS at 5285 Port Royal Road, Springfield, VA 22161.
Copies are available for inspeci orcopying for a fee fun th NRC Public Docmnent Room at 2120 L Street NW. Washington DC-, the PDR g address is Mail Stop .L-6,Washington DC 20555; telephone (202)634-3273, fax (202)634-3343.
1.161-12
16. TJ. Oriesba& and E. Smith, 'A Review ofthe ASME Low Upper Shelf Toughness Evaluation Procedures for Nuclear Reactor Pressure Vessels, NuclearEngineeringandDesign,Volume 130, No. 3, pp. 259-266,1991.
17. Y. Mishima et al., 'Manufacture and Characteristics of a Heavy Section Steel Test Plate with Changing Mechanical Properties in the Though-Thidmess Dircticng NuclearEngineeringandDesign, Volume 137, No. 3, pp. 323-334,
1992.
18. CZ Serpan, Jr., USNRC, Memorandum to C.Y. Cheng, USNRC, "Ratio of Transverse to Longitudinal Orientation Charpy Upper Shelf Energy,' June 25, 1990.4
19. AL KaIlsm, USNRC, Memorandu to C.Y. Chang, USNRC, 'Summay of Fractr Toughness Estimates for Irradiated Yankee Rowe Vessel Materials,' August 30, 1990.4
1.161-13
APPENDIX A:
EXAMPLES
Several cases are provided hem to demonstrate examples of the methods of analysis described in this regulatory guide.
Example 1 (Levels A&B Loading, PWR Vessel)
Consider the following geometric and material properties:
Vessel Geometry and Loading Conditions:
Vessel internal radius, IR= 86.5 in.; A-533B vessel with generic welds Base metal thickness, t = tam = 8.444 in.; Cladding thickness, = 5132 in.
Total thicknes t' - (tm + W - 8.6 in.; Ratio (RN') - 10.06 System accumulation pressure, p. = 2.75 ksi; Cooldown transient = 10 0 Fihr Base Metal Thermo-Elastic Properties:
Modulus of elasticity, E = 27E3 ksi; Poisson's ratio, v = 0.3 Yield stress, a, - 80 ksi; Ultimate stmre, o.= 90 ksi Flow stress, ot= 85 ksi; Fluid heat transfr cocf- =1000 BTU[hr-O-'F
Theral diusivity = 0.98inninte; (E&.,Y(l - v) = 0.305kiPF
Cladding Thermo-Elastic Properties:
Thermal expansion coefficient, a = 9.1E-6PF; Poisson's ratio, v = 0.3 Modulus of elasticity, E - 27E3 ksi; Thermal conductivity = 10 BTUihr-fl-¶F
Stress-free temperature ofcladding = 5500F; Initial operating temp. = 550OF
The VISA-iI code,. with modifications for printing KI, Ka, and KI for 6-to-I aspect ratio flaws, was used to perform analyses for determining transient theamo-mcchanical stresses and temperature gradients across vessel wall thickness. An sxdal flaw with an aspect ratio of 6 to I was postulated to xist in the vessel internal wall. To account for the effect of crack-face pressure on stress intensity factor solutions in VISA-Il, the accumulation pressure was adjusted to be equal to [p.t'.{ l + RPA')}/, 3.02 ksi. At a fixed crack depth of
(0.25t'-O. 1) inch, the tempcrature history prediction is shown in Figre A-I for a transient with a constant cooldown rate of IO0*ihr.
With a factor of safety, SF, of 1.15 on accumulation pressure for Equation I of this guide, the applied J-integral history at a crack depth of (0.25t'+ 0.1) inch for mechanical and thermal stresses, including the cladding effects, is shown in Figure A-2. The applied J-integna reaches the peak steady-state value of 486 in.-Ib/imn in about 150 minutes. Also shown in Figure A-2 are the J-R curves for generic welds (Equations
17, 24-25) at three Charpy V-notch uppcr-sheffencrgy (CVN) values. These J-R curves were drawn for a crack extnsion, As, of 0.1 inch and for the temperature history, in Figure A-I, at a crack depth of
(0.25t'-O.1) inch A study ofFigure A-2 shows an interesting trend that the crack initiation is predicted to take place at about 45 minutes into the transient (with crack-tip temperature of 500F¶) where the applied-J
value (-- 445 in.-lb/i. 2 ) is less than the peak steady-state value and is just equal to the material's J-R curve at CVN value of 40 ft-lb. Thus, the more detailed analysis results in a lower CVN value that satisfies the acceptance criteria.
In order to satisfy Equation 2, with a safety factor of 1.25 on accumulation pressure, Figure A-3 shows that CVN value should be greater than or equal to 41 ft-lb. This is significantly lower than the 47 fi-lb value obtained by using the steady-state applied J-integral approach for analyzing transients with constant cooldown rates.
'F.A Simoaen gt at, VISA-HI - A Computer Code for Preddting the Probabfity oafReacdr Presmer Vene F&Rlhe, USNRC.
NUREOICR-4436, March 1926.
A-1
Example 2 (Levels C and D Loading, PWR Vessel)
The problem statement was presented in a meeting of the ASME Section XI Working Groups on Flaw Evaluation and Operating Plant Criteria (in Louisville, Kentucky, on December 1, 1992), where results of the analyses were compared by the participants. The vessel geometry and material properties are:
PWR vessel internal radius, R4 - 90.0 inch; A-533B plate material thickness, t = tsu - 9.0 inch; Cladding thickness, tI, = 0, Rf& = 10 Copper, Cu = 0.35wt%; Nickel, Ni = 0.3 wt%; Initial RTm. = 0.01F
Pre-irradiated CVN, - 108 R-lb (L-T orientation)
Surface fluence, *t = 3.0E19 n/cm2 Flaw orientation = Axial, in plate material; Flaw aspect ratio 6 to I
Fluid temperature at vessel surface, T(tm) - [550 - 25011 - exmg- 0.1 tm)W]'F with time, tin, in minutes.
Heat transfer coefE = 320 BTU/br-t-OF; Thermal diffusivity = 0.98 in. 2/min Elastic modulus, E = 28E3 ksi; Poisson's ratio, v =0.3; a f S&IE-6 inlm.-0F
Yield stress, oy = 80 ksi; Flow stres, of-85 ksi J-R curve: J = (SF).[CI.(Aa)c. cxpfC3.(Aa)"4] in.-kipr/ 2 where.
ln(Cl) = [-2.89+1.22 ln(CVN,) - 0.0027 T + 0.014 ((t)]
C2 = [0.077 + 0.116 ln(Cl)]
C3 = [- 0.0812 - 0.0092 ln(Cl)]
C4 -0.417 SF = 0.741 forLevelC events The VISA-i code was used to determine thermal stress and temperature history for the Level C
transient specified in the problem. It was found that at time tm = 20 minutes, the peak thermal stresses occur.
The corresponding peak thermal strews intensity factor as a function of crack depth to vessel thickness ratio, a/, of semi-elliptical flaws is given as:
Kj- [21.026+374.22(a/t)-1593.56(aA)0+2912.1(a/t)-2029.7(a/)] ksih"m. with 0.05 ! at
- 0.5 Therefore, at a = I inch, K, - 46.6 ksi-"in. At an internalpressure, p - I ksi, the pressure induced K, - 18.9 ksi-fin. Now, ifthe pressure, p, is increased, then at a pressure of 6.75 ksi, the J-applied at a = (0.It + t. +
0.1) inch becomes equal to the material's J-R curve as shown in Figure A-4. This will mark an "initiation" of ductile flaw growth. The temperature at the crack-tip (a= 0.It + t,) for time tmn= 20 minutes is 400OF. If internal pressure p is further increased, in Figure A-4 it can be seen that at pressure p = 7.56 ksi the crack growth becomes unstable. That is, the slope of the J-applied curve becomes greater than the slope of the material's J-R curve.
Example 3 (Levels C and D Loading, BWR Vessel)
The problem statement is the same as in Example 2, except for a BWR vessel geometlry. The vessel geometric details are:
BWR vessel internal radius, R.= 120.0 inch; A-533B plate material Thickness, t = t, - 6.0 in.; Cladding thickness, t., = 0, M - 20
Flaw orientation - Axial, in plate material; Flaw aspect ratio - 6 to I
The VISA-il code was used to determine thermal stress and temperature history for the Level C
transient specified in the problem. It was found that at time tin =16 minutes, peak thermal stresses occur.
The corresponding peak thermal stress intensity factor as a function of crack depth to vessel thickness ratio, alt, of semi-elliptical flaws is given as:
KI,= [I 2.243+227.94(at)-972.71 (aht)4+1785.2(a/t)3 -1249.3(at)4] ksi"in., with 0.05 < a/t < 0.5 A-2
Therefore, at a = I inch, K -=27.9 ksi-"in. At an internal pressure, p = I ksi, the pressure-induced K -=37.0
ksiWin. If the pressure, p, is increased, at a pressur of4.55 ksi, the J-applied at a = (0.1t + tL +O. I) inch becomes equal to the material's J-R curve as shown in Figure A-5, which will mark an "initiation" of ductile flaw growh The temperature at the crack tip (a-= 0.1t + t.) for time tmi= 16 minutes is 4050F. If the pressure, p, is further increased (see Figure A-5), it can be seen that at a pressure p = 4.75 ksi the crack growth has become unstable. The slope of the J-applied curve is now greater than the slope of the materials J-R curve.
Example 4 (Thermal K1t for Prescribed Levels C and D Leading, PWR Vessel)
For a PWR vessel, thermal K3 values are determined for a few prescribed cooldown rate (CR)
transients. The geometric and material properties are iven as:
Vessel Geometry and Loading Conditions:
Vessel internal radius, Rý = 86.875 in.; A-533B plate material with cladding Base metal thickness, t = tm - 8.625 in.; Cladding thickness, t. = 5/16 in.
Total thickness, t' = (thj + W -98.9375 in.; Ratio, W/t') = 9.72 Thermal cooldown rate, CR = 100"Fihr to 600"Fihr (constant, for each analysis)
Inner wall temperature, Tk .(R - = 550'F; TsJR = R) = 150F
Base Metal Thermo-Elastic Properties:
Modulus of elasticity, E = 27E3 ksi; Poisson's ratio, v = 0.3 Fluid-film heat transfer coefficient,= 1000 BTUihr-f -PF
Thermal diffusivity = 0.98 in!/minute; (Ea)/(I - v) = 0.305 Cladding Thermo-Elastic Properties:
Thermal expansion coefficient, a = 9.1E-61F; Poisson's ratio, v = 0.3 Modulus of elasticity, E = 27E3 ksi; Thermal conductivity- 10 BTU/hr-ft-OF
Sftr -free temperature of cladding = 550-F; Initial operating temp. = 550WF
The VISA-i1 code was used to determine temperature and thermal stress history for constant CR transients of
0
1 0F/hr, 15( 0F/hr, 20(0Fir, 30(0F/hr, 400-F/hr, 500rFihr, and 600¶bhr. The corresponding peak thermal stress intensity factors, Y41, as a function ofcrack depth to vessel thickness ratio, aWt,, for 6-to. i aspect ratio semi-elliptical flaws, were computed using the VISA-11 code. These are shown in Figure A-6 and are presented here in polynomial expressions using least-square fits as:
For CR - 100"Fhr with 0.05 <(aft') < 0.5:
Ka - [27.284 - 5.838 (aft') - 0.3548 (aft') 2 - 8.3858 (aht'Y] ksifin.
For CR - 150°F/hr, with 0.05 g (aft') s 0.5:
K,4 = [32.003 + 40.012 (aft') - 138.2 (a/t'r - 113.98 (aht')r] ksif"in For CR = 200F/hr with 0.05 < (aft') g 0.5:
KI = [36.362 + 82.011 (aft') - 265.01 (aft')r + 226.9 (aft')2] ksi/rin.
For CR - 300F7hr with 0.05 : (aft') 10.5:
K,4 = [43.667 + 150.77 (aft') - 474.9 (aft')2 + 415.01 (a/t')2] ksijrin.
For CR - 400°F/hr with 0.05 r. (aft') < 0.5:
Kft = [49.254 + 201.12 (aft') - 632.1 (a/t'r + 557.87 (a/t')3] ksii"in.
For CR - 500OF/hr with 0.05 < (aft') < 0.5:
4 = [53.552 + 237.64 (aft') - 749.6 (aA')2 + 666.62 (aft') 3] ksarin For CR - 600F*hr with 0.05 g (at') < 0.5:
K,, = [56.927 + 264.21 (a/t') - 838.6 (aft')2 + 750.88 (aft')2] ksiv"in.
A-3
These results were also used in developing the unified Equation 13 for K, where the constant CR and the nomalizhd crack depth, at', are used as dependent vaiables. A least-square statistical fit was performed to obtain Equation 13. The czoss-product term, (CRXa/'), was also used in developing this fit, in addition to the polynomial terms in aW and CIL
A-4
CD-
t"o L.
0
iz O0
4- an PWR Vessel, Service Levels A and 8
11R=86.5 In, t =0.156 In, t'8.6 I,. a/t*=0.25 l O0"/hr Cooldown for 150 Minutes
0 *o--o Temperature at the Crack-Tip
30 ..... 60 ..... 9 ..... 0 ""150
Transient Time (Minutes)
Figure A-i: Transient Tcmperatxnc H-istory at Crack Tip for Scrvce Levels A and B.
A-5
0 -.
Co
0
U)
.-D1 PWR, Level A&B, Generic Welds. t.L0. 156 In C R3,1486.5in, t'=8.6 In, Axial Crack, a/V'0.25 Io00*F/hr Cooldown, Critical Time =45 Minutes o---o : J (Applied Load, Criteria No. 1)
Crack Extension = 0.1 In.
+-+ : J-R Curve (CVN - 41 ft-1b)
O-'-- : J-R Curve (CVN - 40 ff-ib)
- " * J-R Curve (CVN = 39 ft-lb)
6 . 3b 60 90 120 150
Transient Time (Minutes)
Figure A-2: J-Applied mnd J-R Curve History at Crack Tip for Service Levels A and B at a Crack Extensio of 0.1 inch.
A-6
0
0
C
N
CE)
-o T
C
4 C
o PWR. Level A&B, Generic Welds, f.O=.l 56 In j R=86.51n, f=8.6 In.. Axial Crack, a/t=0.25 S
o00F/hr Cooldown. Critical Time =45 Minutes o--o : J (pplied Load. Criteria No. 1)
ICe---1.. J-%Curve (CVt=41 ft-lb, T=5.00TF)
I
..
,--
...
.J (Applied Load, Criteria No. 2)
.......... ,.,,.
i .
l
.0 "0.1" 0.2 0.4 0.5 0.6 Crack Extension (in)
Figure A-3: Acceptable Upper-Shelf Ewrgy in a PWR Vessel for Service Levels A and B.
A-7
,I ; i I
0
v V
Un C4 K
0
0
CdA
.0
T
-o
- PWR Vessel, Level C, High Toughness Plate R1=90 In, t=9 In, Axial Crack, ao0.9 In T=550-250[exp(-.1 tm)]°F, tm=20 minutes
- J (Appliled Load, pi = 6.75 ksl)
"1--43 :J-R Curve (CVNP= 108 ft-lb, T=4000F)
-- : J (Applied Load, ph = 7.56 ksl) K
i'5 0 a a 0.1 0.2 0.3 0.4 0.5 0.6 Crack Extension (in)
Figure A4: Safdy Margin Evaluation in a PWR Vessel for Service Lcve C.
A-S
0
0.2 0.3 0.4 Crack Extension (in)
Figure A-5: Safety Margin Evaluation in a BWR Vessel for Sevice Level C.
A-9
I
- I.- I8 .
td.'Ca 5/16
8.93 Ii.
a..
(-60F0r.
F-50
M.
+.4 r C
K
I-V).
X Cý <-00 FlIV.
<-0 Fltw.
0 . . . , , ,, , , , , , ,, , , , , Tftnllfly -IseF
0 0.2 0.3 0.4 0.5 I
Normalized Crack Depth (a/f)
Figure A-6: Peak Thermal Stress Intensity Factors in a PWR Vessel for Transients with Several Different, but Constant. Cooldown Rates.
A-1O
"APPENDIXB
COMPUTATION OF STRESS INTENSITY FACTORS
Information about computing transient temperature gradient across the vessel wall thickness, themal stresses, pressure, and thermal stress intensity factors (K.,, Kt) are provided in this Appendix as FORTRAN
subroutines fiom the VISA-1 code. Additional details on the computational method, theory used, limitations, and names of the major variables used are available in NUREG/CR-4486' and NUREG/CR-3384.' The computer code provided in this Appendix is for general illustration only, to show how the cladding effects could be incorporated for thermal stresses and thermal sres intensity factors caused by differential fthemal expansion between the cladding and the base metal. i.censees should ensure that the computer codes they use include an idepth evaluation of ts effiects.
A description of cladding-iduced thermal stress intensity factors is presented in Appendix A to NUREOICR-4486. Limitations of the stress intensity factor correction factors for finite length semi-elliptical surface flaws are indicated in Appendix C to NUREG/CR-4486. In developing these correction factors, only uniform membrane and linear bending stresses were considered. In addition, the correction factors for circunfe-ential flaws were assumed to be the same as the ones for axial flaws. Improved solutions may be used on a case-by-case basis ifjustified.
'F.A Shnonen et al, *VISA-H - A Computer Code for Predicting te Probability of Reactor Pressure Vessel Failure* USNRC,
NUREG/CR-4486, March 1986. D.L Stevens et al., 'VISA- ACoemputer Code for Predicting the hbability of Reactor Pressure Vessel Failure," USNRC, NUREO/CR-33S4, September 1983. Copies are available for inspection orcopying fora fee from the NRC
Public Document Roomn at 2120 L Stee NW., Washingto DC, the PDR's ai'ling address is Mail Stop U,6. Washingtm, DC
20555; telephone (202)634-3273; fax (202)634-3343. Copies ofNUREOICRs may be purchased at curent rates from de U.S.
Government Printing Office, Post Office Box 37082, Washington, DC 20013-7082 (telephone (202)512-1200); or from the Natioad Tedmical Information Service by writing NTIS at 5285 Poat Royal Road, pringd, VA 22161.
13- I I
Taken From: VISA-H Code LNUREG/CR-4486 (1986). NUREQXR-3384 (1983)]
SUBROUTINE SPKI
Calculate PresstireValuesand, Stren Intensity Factor, PKI
DIMENSION CONST(5)
REAL 1(5). IC(5)
INTEGER CRAM TBE
C DETERMNE POLYNOMIAL REPRESENTATION OF PRESSURE
CONST(l) = PDATA(I)
CONST(2) = ((-2.5)*PDATA(I)+48*PDATA(2)-36*PDATA(3)+
I 16*PDATA(4)-3*PDATA(S))1(3*TMAX)
CONST(3) = (35*PDATA(l)-104*PDATA(2)+114*PDATA(3)
I 56*PDATA(4)fI I *PDATA(5))*V(3*TMX**2)
CONST(4) = ((-5)*PDATA(])+18*PDATAC2)-24*PDATA(3)+
I 14*PDATA(4)-3*PDATA(5))*161(3*TMAX**3)
CONST(5) = (PDATA(l)-4*PDATA(2)+6*PDATA(3)-4*PDATA(4)+
I PDATA(5))*32/(3*TMAX**4)
C Calculate PRESSURE Component of Applied K, PKI, For Each 1-kne Crack Depth OUTRAD=RAD+TH
FACTOR = RAD**2.0 / (OUTRAD**2.0 - RAD**2.0)
C
DO 120 TAM = 1, 10
IT = TMAX*TflvWI0.0
DOI 10 CRACK= 1, 1CMAX
X=Z(CRACKYM
C CALCULATE INFLUENCE COEFFICIENTS
DO 100M= 1,5 IM =ZZ(Kl)+X*ZZOA.2)+(X**2)*ZZ(K3)+(X**3)*72XK4)
IC(M) - ZZC(MI) + X*ZZC(K2) + (X**2)*ZZC043) + (X**3)*ZZC(K4)
100 CONTINU13 PRES(TDAE) = CONST(I)+CONST(2)*TT+CONST(3)*TT**2+CONST(4)*TT*
I *3+CONST(5)*TT**4 PKI(CRACK.TIME) = PRESCrB4E)*((3.1416*Z(CRACK))**.5)*(10.5238*1(1)
I -1.1524*I(-2)*X40.1729*1(3)*(X**2)-0.0230*1(4)
2 *(X**3)+0.0029*1(5)*(X**4))
B-2
PKIC(CRACK,TIMfE) = 5*pREs(TIME)((3. 1416*Z(CRACK))**.5)*1C(I)
RATIO = RAD / (1O.0*5 TH)
PKI(CRACK.TIME) - RATIO
- PKI(CRACKTIME)
PKIC(CRACK,TIME) = RATIO
- PKIC(CRACK.TIME)
c CALCULATE HOOP STRESS
SHOOP(CRACKTIE) = FACTORS*PRES(TME)
1 (1.0 + (OUTRAD/CRAD + Z(CRACK)))**2.0)
110 CONTINUE
C CALCULATE LONGITUDINAL STRESS
SLONCTCITIME) = PRES(TIME) *FACTOR
120 CONTINUE
RETUR}N
END
SUBROUTINE TPOLY
C CALCULATE WATER TEMPERATURES USING A POLYNOMIAL' MODEL
REAL TEMP(5), CONST(5), S(5). AN(4). Y(4,5). KTEST
REAL K, KO, CP(4), SUM(4)
INTEGER TIME, CRACK. CONSTK CONSTE
INTEGER Q
C KPOLYNOMIAL Modeling of The Wate Temperature C Determine Meta Temperature For EACH CRACK DEPTH AND TIME INTERVAL
DO 100N= 1,5 TEMP(N) - TDATA(N) - TINT
100 CONTINUE
C FIT A 'POLYNOMIAL TO THE WATER TEMERATURE
CONST(1) = TEMP(1)
CONsT(2) - ((..25)*TEMP(1) + 48*TEMP(2) - 36*TEMP(3) +
1 16*TEMP(4) - 3*TEM[P(5WY(3*TMAI)
CONST(3) - (35*TEMOP(1) - 1O4*TEMP(2) +11 4*TEMP(3)
I 56*TEMP(4) + II *TEhe(5))*2/(3*TMAX**2)
CONST(4) = ((-5)*TEMP() + 18*TEMP(2) -24*TEMP(3) +
1 14*TEMP(4) - 3*TEMP(5)Y'16/(3*TMAX**3)
CONST(5) = (TEMP(1) - 4*TEMP(2) + 6*TEMP(3) - 4*TEMP(4) +
I TEIP(5))*32I(3*TMAX**4)
DO 150 TIME =1. 10
B-3
"TT = TMAX*TIMFJIO.
C EQUATION FOR THE TEMPERATURE OF THE WATER
TWATER(TIME) = TINT+CONST(I)+ CONST(2)*TT + CONST(3)*TT**2 +
1 CONST(4)*TT**3 + CONST(5)*TT**4 DO 150 CRACK= 1,5 K=KO
I1O X = ZQ(CRACK)/TM
TAU - K*TTfTH**2 DO 120M= 1, 5 W(M) = CONST(M) *
120 CONTINUE
DO 130N- 1,4 ALNQ - AL(NQ)
AN(N) = 2
- SIN(ALNQ)/(ALNQ + SIN(ALNQ)* COS(ALNQ))
CP(N) = COS(ALNQ (I-X))
Y(N,I) -I - EXP(-(ALNQ**2)*TAU)
DO 130 M -2, 5 Y(NM) = TAU**(M-I) - (Y (NM-IyALNQ**2)*(M-1)
130 CONTINUE
DO 140N= 1,4 ,
ALNQ - AL(NQ)
SUM(N) - AN(N)
- CP(N) * (S(I)
- EXP(-(ALNQ**2*TAU)) + S(2)
I
- Y(N,IYALNQ**2 +2*S(3)* Y(N,2)ALNQ**2 + 3 *S(4)
- Y(N,3)
2 /ALNQ**2 +4 *S(5)*Y(N,4)/ALNQ**2)
140 CONTINUE
C EQUATION FOR THE QUARTER POINT TEMPERATURES
TQ(CRACK,TIME) - TWATER(TIME) - SUM(I) - SUM(2) - SUM(3) - SUM(4)
C CONTROL FOR THE CONSTANT KAPPA OPTION
IF (CONSTK-.EQ. 1)GO TO 150
C TESTFOR THE ACCURACY OF KAPPA FOR THE GIVEN METAL TEMPERATURE,
C IF THE DESIRED ACCURACY IS NOT OBTAINED, ITERATE ON KAPPA
C FOR THIS CRACK DEPTH AND TIME.
KTEST = 1.030 - (5.97E-7)*((T(CRACKTIME))**2)
IF ((ABS(KTEST-K)) JLE. 0.0001) GO TO 150
K=KIEST
GOTO 110
150 CONTINUE
B-4
RETURN
END
SUBROUTINE TEXP
C Calculate WATER TEMPERATURES Using an 'Exponential Decay" Model REAL B. KTEST, K, KO. SUM(4)
INTEGER CRACK, TIME, CONSTK, CONSTE
INTEGER Q
C EXPONENTIAL DECAY MODEL OF THE WATER TEMPERATURE
DO 130 TIME = 1, 10
TT = TMAX*TIMFl0.
C EQUATION FOR THE TEMPERATURE OF WATER
TWATER(TIME) = TO + DT * (I-EXP(-BE*TT))
DO 130 CRACK i 1,5 K=KO
100 WSQ = BE*TH*THK
TAU - K*TT/(TH*TH)
DO 120N=1,4 ALNQ = AL(N.Q)
B = -DT*((2*SIN(ALNQ)/(ALNQ+(SIN(ALNQ))*(COS(ALNQ))))
I *(EXP(-(ALNQ**2*TAU))-EXP(-WSQ*TAU))I((ALNQ**2/WSQ)-1))
X =i ZQ(CRACKYTH
SUM(N) = B
- COS(ALNQ*(I -X))
120 CONTINUE
C EQUATON FOR THE -QUARTER POINTS" TEMPERATURE VALUES
TQ(CRACKTIME) = TWATER(TIME) - SUM(1) - SUM(2) - SUM(3) - SUM(4)
C CONTROL FOR THE CONSTANT KAPPA OPTION
IF (CONSTK.EQ. 1) GO TO 130
C TEST FOR KAPPA ACCURACY AND CONTROL OF KAPPA OPTION
KTEST = 1.030 - (5.97E-7)*((T(CRACKTIvE))**2)
IF ((ABS(KTEST-K)) IE. 0.0001) GO TO 130
K =KTEST
GO TO 100
130 CONTINUE
RETURN
END
B-5
,I i ,
SUBROUTINE SKIT
K
C Calculate Stress and Temperature at Crack-Tip and Thermal Stress C Intensity Factor, SKIt REAL E(5,10), CC(5), I(5), IC(5)
INTEGER CRACK, TIME
INTEGER Q, CONSTE, CONSTK
C DETERMIINE POLYNOMIAL REPRESENTATION OF TEMPERATURE PROFILE
C CONVERT CLAD TIHERMAL CONDUCTIVITY TO INCH AND MINUTE UNITS
CCOND = CCOND / (12.0*60.0)
COND = COND /(12.0*60.0)
DO 105 TIME - 1, 10
TQI = TQ(I.TIME)
TQ2 - TQ(2,TIME)
TQ3 - TQ(3,TIME)
TQ4 = TQ(4.TIME)
TQ5 = TQ(5,TIME)
Cl =TQI
C2 = (-25*TQI+48*TQ2-36*TQ3+16*TQ4-3*TQS)/(3*TH) K
C3 = (35*TQI-104*TQ2+114*TQ3-56*TQ4+1 I *TQS)*(2.0/3.0*TH**(-2))
C4 = (-5*TQI+1 S*TQ2-24*TQ3+14*TQ4-3*TQS)*(16.03.0*TH**(-3))
C5 = (TQI-4*TQ2+6*TQ3-4*TQ4+TQ5)*(32.0/3.0*TH**(-4))
C CALCUATE TEMPRATURE AT THE CRACK TIPS
DO 100 CRACK = 1, ICMAX
T(CRACKTIME) = C1+C2*Z(CRACK)+C3*(Z(CRACK)**2)
I -C4*(Z(CRACK)**3)+C5*(Z(CRACK)**4)
100 CONTINUE
IF (CTH .LE. 0.0) GO TO 105 T(,TIME) = T(2,TIME) - (COND/CCOND)*(r(2,TIME)-T(ITME))
105 CONTINUE
IF (CONSTE .EQ. 1) GO TO 120
DO 10 TIME = 1, 10
DO 110 CRACK = 1,5 E(CRACKTIME) = 0.286+(5.400E-5 * (TQ(CRACK,TIMED)))
1 -(2.600E-8 * (TQ(CRACK,TIME))**2)
110 CONTINUE
B-6
GO TO 140
120 DO 130 TIME = 1.10
DO 130 CRACK - 1,5 E(CRACK,TIME) - EDATA
130 CONTINUE
C DETERMIN POLYNOMIAL REPRESENTATION OF STRESS DIST
140 DO 170 TIME - 1, 10
DO 150 CRACK - 1, 5 CC(CRACK) - E(CRACKTIMAE)*TQ(CRACK.TIIE
150 CONTINUE
Al =CCWI
A2 - (-25*CC(1)48*CC(2)-36*CC(3)+16*CC(4)-3*CC(5)Yt3.0
A3-(35*CC(1)d104*CCC)+1 14*CC(3)-56*CC(4)+1 1*CC(5))*(2.0/3 0)
A4 = (-5*CC(1)+18*CC(2)-24*CC(3)+14*CC(4)-3*CC(5))*(16.0I3.0)
A5 - (CC(l)-4*CC(2)+6*CC(3)-4*CC(4)+CC(5))*(32.0I3.0)
SIG) = AMf.0 + A3/3.0 + A414.0 + A5/5.0
SIG2 =-A2 SIG3 = -A3 SIG4 = -A4 IGS = -A5 C CALCULATE STRESS AT CRACK TIPS
DO 170 CRACK -
1. ICMAX
X =Z(CRACK)/TH
STRESS(CRACK,TIME) - SlGI + S102*X + SIG3*(X**2)
I + SIG4*(X**3) + SIGS*(X**4)
C CALCULATE INFLUENCE FUNCTIONS
DO 160M= 1,5 I(M = ZZ(M1) +X*ZZ(M,2)+ (X**2)*ZZQM.43)+ (X**3)*ZZ(M.4)
IC(M) =ZZC(Mj)+X*ZZC(M2)4(X**2)*ZZC(MK3)+(X**3)*ZZC(MA)
160 CONTINUE
A = Z(CRACK)
C EQUATION FOR THE THERMAL STRESS INTENSITY
TK(CRACKTIE) - ((3.1416*A)**.5)*(SIGI 1(1)
I +51G2*I(2)*X+SIG3*I(3)*X**2
2 +SIG4*IK4)*X**3+SI050I(5)*X**4)
TKC(CRACK,TilvE) =(3. 141 6A)**.5)*(SIO1 *IC(1)+SIG2*IC(2)
I *X+51G3*IC(3)*X**2+5104*IC(4)*X**3+5105*IC(5)*X**4)
B-7
. I I I
170 CONTINUE
RETURN
END
SUBROUTINE KICLAD
C THIS SUBROUTINE CALCULATES STRESSES AND STRESS INTENSITY FACTORS
C DUE TO THE PRESENCE OF "CLADDING" ON THE I.D. SURFACE OF THE VESSEL
INTEGER CRACK, TIME
INTEGER CONSTE, CONSTK, Q
REAL 10, 11 DO 170 TIME- I, 10
C CALCULATE STRESS DISTRIBUTION THROUGH VESSEL WALL
C TEMP AT CLAD/BASE METAL INTERFACE
TI = 0.5*(T(2,TIME) + T(3,TIME))
C TEMPERATURE AT THE VESSEL I.D.
TO - T(,TIME)
C STRESS-FREE TEMPERATURE
TI= SFREET
C CALCULATE STRESS DISTRIBUTION DUE TO CLAD K
C SIGCI = STRESS IN CLAD AT VESSEL I.D.
C SIGC2 = STRESS IN CLAD AT CLAD/BASE METAL INTERFACE
C SIGBI = STRESS IN BASE METAL AT CLAD/BASE METAL INTERFACE
C SIGB2 - STRESS IN BASE METAL AT VESSEL O.D.
DELEA - CLADE*CALPHA*(I-ARATIO)/(I-CLADNU)
C CALCULATE STRESS IN CLAD (KSI)
SIGCI - DELEA* (TI - TO)
SIGC2 = DELEA ( - TI)
C CALCULATE FORCE DEVELOPED IN CLAD
FCLAD = CTH*0.5*(SIGCI + SIGC2)
C CALCULATE STRESSES IN BASE METAL (KSI)
RO -RAD
RI = RAD +CTH
R2 =RAD+TH
CONST = 1.0/((R2/RI) 2.0-1.0)*(RO-RI)/RI *DELEA
I *(TI-O.5*(TO+TI))
SIGBI = CONST * (I + (R2/RI)**2.0)
B-8
SIGB2 - CONST
- 2.0
C CALCULATE FORCE DEVELOPED IN BASE M[ETAL
FBASE - (CTH-TH*0.S*(SIGBI+SIGB2)
C ADJUST SIGB 1 AND'SIGB2 TO BALANCE FORCES FCLAD AND FBASE
SIGINC - 0.5*(SIGBI-SIGB2)
SIGAVE -0.5*(SIGBI+SIGB2)*FCLAD/FBASE
SlOB 1 - SIGAVE + SIGINC
SIGB2 - SIGAVE - SIGINC
C CALCULATE CONSTANTS DESCRIB3ING STRESS DISTRIB3UTION
C QI - SLOPE OF CLAD STRESS DISTR.
QI - (SIGCI-SIGC2)/SIGC1/(CTHITH
C P -SLOPE OF BASE METAL STRESS DISTR.
P = (SIGB2-SIGB1Y/SIGCJC/ I((IH-CTHYMH
C -R -INTERCEPT OF BASE METAL STRESS GRAD. AT VESSEL L.D.
R-=-(SIOBMI/SIO - P*CTWFH)
C CALCULATE STRESS AND KI DUE TO CLAD FOR ALL Z(CRACK'S,
C KI AT THE L.D. SURFACE EQUALS ZERO (LE.,CRACKDEPTH = ZERO)
SCLAD(1,TJME) - SIOCI
CLADK(1,TIME) = 0.0
C KIlIN CLAD NEAR CLAD/BASE METAL INTERFACE
SCLAD(2,TDME) = SIGC2 ALP = Z(2)/TH
10 = 1.l22+0.9513*ALP-0.624*ALP**2.0+8.3306*ALP**3.0
Il = 0.6825+0.3704*ALP-O.0832*ALP**2.O+2.8251 *ALP* $3.0
CLADK(2,TIME) - SQRT(3. 14159*Z(2))*SIGC1 *(IO-QI*ALP*I1)
C CALCULATE KI IN BASE M[ETAL
XI= CTH/TH
DO 170 CRACK = 3,35 ALP - Z(CRACK)/T
SCLAD(CRACKTIME) = (-R+ALP*P)*SIGCI
10 - 1.12240.9513*ALP-0.624*ALP**2.0+8.3306*ALP**3.0
CLADK(CRACK,TIMvE) - SQRT(3. 14159*Z(CRACK))*SIGC1 *1.751938 I *((IO-0.63662)*((1 .0+R)*ASINQUI/ALP)+ALP*((QI+R*P)
2 *SQRT(1 QWXALP)**2.)-QI)-1 .570796*R)+(IO01 .0)*(((1 .0+R)-XYt.
3 *(QI+R*P))*SQRT(1 .- (XALP)**2.)+ALPt2.0*(QI+R*P)*ASIN(XIIALP)
4 -1.0-0.7894*R*P*ALP))
170 CONTINUE
B-9 I I
RETURN
END
SUBROUTINE FACMB (AAA, BBB, THH, FMA, FMB, FBA, FBB)
C THIS SUBROUTINE CORRECTS FOR *FINITE LENGTH" SEMI-ELLIPTICAL FLAWS
DIMENSION ZM(2,4), ZB(2,4), Z(2)
DIMENSION X1(12), YM(12,4), YB(12,4), Y(4)
DATA XI/O., .0125, .025, .0375, .05, .075,.l,.15,.2,.3,.4,5I
DATA Y/.05, .25, .5, .8 /
DATA YM/ 1.0,.99,.98,.96,.95,.91,.87,.80,.75,.66,.60,.55,
1 1.0,.94,.88,.83,.80,.76,.73,.68,.63,.55,49,.44,
2 1.0,.88,.77,.69,.64,.59,.55,.49,.44,36,31,.27,
3 1.0,.72,.56,.48,.43,.38,.35,.29,.24,. 18,.15,.131 DATA YB/ i.0,98,97,.95,.94,.92,.89,85,.82,74,.66,.58,
2 1., .93,.88,.84,.80,.75,.72,.67,.63,.57,.50,.43,
2 1., .84,.71 ,.63,.57,.49,.45,.39,.35,.29,.23,.18,
3 1., .69,.50,.38,.29,.20,.14,.08,.05,.02,-.01,-.041 DATA Z/ 0.0, 0.5 /
DATA ZM/.44,.55, .40,.48,.31,.31,.23,.17 I
DATA ZB/.50, .62, .63, .67, .58, .50, .43,.32 I
AOL = AAAI(2.0*BBB)
AOT = AAAfTHH
DO 100- I=1,3 J=I
IF(Y(I+1).GT.AOT) GO TO 110
100 CONTINUE
liONI =J
N2 =J+l DO 1201= 1, 11 J=I
IF(XI(I+1).GT.AOL) GOTO 130
120 CONTINUE
130 MI =J
M2 = J+l FACI = (AOL-XI(M1))/(Xl(M2)-XI(M1))
XXI = YM(MI ,N1)+FAC I *(YM(M2,NI)-YM(MI,N1))
B-10
MU =YM(MIN2)+FACI*(YM(M2,N2)-YM(MI.N2))
FAC =(AOT-Y(NI)Y(Y(N2)-Y(Nl))
FMA = XXI + FAC*(XX2 - XXI
XXI = YB(MINl) + FACI*(YB(tMNl)-YB(MINl))
xm -YB(miN2)+FACI*(YB(tMN2)-YB(MI.N2))
IFBA - XXI + FAC*(XM - XXI
FAC I = AOL/0-5 XXI -ZKI.Nl)+FACI*(U4C2,Nl)-ZM(INl))
)m =zm(iN2)+FACI*(ZM(2,N2)-ZM(l.N2))
FMB = XXI + FAC*(3M-XXI)
XXI - ZB(INl) + FACI*(ZB(2,Nl)- ZB(INl))
XM = ZB(IN2) +FACI*(ZB<2.N2)- ZB(I.N2))
FBB - XXI + FAC*(XM - XXI)
RETURN
END
B-11
REGULATORY ANALYSIS
1. STATEMENT OF THE PROBLEM
Appendix 0, 'Fracur Toughness Requircmencs,' to 10 CFR Part 50, 'Domestic Licensing of Production and Utilization Facilities,' requires, in part, that the reactor vessel bcltline materials '... must have Charpy upper-shelf energy of no less than 75 ft-lb (102J) initia"ly and must maintain uppcer-shelf energy throughout the life of the vessel of no less than
50 ft-lb (68J). unless it is demonstrated in a manner approved by the Director, Office of Nuclear Reactor Regulation, that lower values of upper-shcfenergy will provide margins of safety against fracture equivalent to those required by Appendix ofothe ASME Code." This Regulatoxy Guide 1.161, 'Evaluation of Reactor Prcssure Vessels with Charpy Upper-Shelf Ener-y Less Than 50 ft-lb,' has been developed to provide acceptance criteria and analysis methods acceptable to the NRC
staff for demonstrating margins equivalent to those in Appendix 0 to Section III of the ASME Code.
Publication ofrgulatcry guidance was undertaken because no comprehensive guidance currently exists, and there are reactors, both pressurized water reactors and boiling water reactors, with upper-shelf energy that is projected to fall below the 50 ft-lb regulatory limit before the end of the current license period. Withut comprehensive regulatory guidance. each affected licensee will have to submit a plant-specific analysis, including acceptance criteria and evaluation methods, and the staffwill have to evaluate each submittal without the benefit of staed acceptance criteria and approved evaluation methods.
2. OBJECTIVES
The objective of this guide is to provide acceptance criteria and evaluation methods acceptable to the NRC staff for demonstrating margins equivalent to those in Appendix 0 to Section MI of the ASME Code for those beltline materials whose Charpy upper-shelf energy falls below the regulatory limit provided in Appendix 0 to 10 CFR Part 50.
3. ALTERNATIVES
Two alternatives to issuing evaluation procedures for pressure vessels with Charpy upper-shelf energy less than 50
ft-lb were considered: (1) endorse actions being implemented by Section X[ of the ASME Code and (2) take no action.
3.1 Endone ASME Code,Section XI, Appendix K
The ASME, in Section X)M has published Appendix K' that provides acceptance criteria and evaluation procedures for pressure vessels with Charpy upper-shclf energy less than 50 ft-lb. However, the Appendix K evaluation procedures curently address only Service Levels A and B, and no guidance on specific materials properties is provided. It is important that all four service levels be considered in the evaluations, and it is important that specific guidance on estimating material properties be provided. Given the ASME codification process, and the process whereby the NRC endorses ASME
appendices and code cases, the time delay in obtaining suitable guidance would be excessive. At present, the ASME's Appendix K does not provide complete guidance. As discussed above, Appendix K does not provide information on the selection of transients, and it gives very little detail on the selection of material properties. As such, a request for revision of Appendix K to Section XM of the ASME Code will have to be made.
3.2 Take No Action As discussed in SECY-93-048, 2 'Status of Reactor Pressure Vessel Issues Including Compliance With 10 CFR Part 50, Appendices 0 and IV,using the NRC staffs generic criteria for estimating Charpy upper-shelf energy, there are currently
15 plants that would have calculated upper-shelf energy less than 50 ft-lb and 3 others that would haveuppcr-shelf energy below 50 ft-lb before the end of their operating licenses. Appendix 0 to 10 CFR Part 50 requires that licensees submit
,A.npdxK emously, Code Caw N-*S 12), "Asesment cef tor Vessels with low Upper ShelfCmpy kmpact Er"y ls,' Amrica Society of Mechanical Engineers, Section Xn 1993.
2 Jcmes MKTaylor, Executive Director for Opcrat, SECY-93-04, Policy Issue (Infamtloo) for the Connissionae, USNRC Fchuay 25, 199.
ava'lable fr"orm ics am~ eorpying f a fee =fr9= NRC Public Document Room at 2120 L Sret NW., Washington, DC; 1he PDR's maiing address is Mail Stop LL.6, Washgto. DC 20555; telephone (202)634-3273; fax (202)634-3343.
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analyses to demonstrate margins equivalent to those in Appendix G to Section MI of the ASME Code 3 years before the upper-shelf energy of any beitline materials falls below 50 ft-lb. Therefore, taking no action is not a viable alternative.
4. COSTS AND BENEFITS OF ALTERNATIVES
The cost and benefits of the two alternatives discussed above are presented here.
4.1 Endorse Appendix K to ASME Code Section XI
The acceptance criteria proposed in Appendix K to ASME Section XI are identical to those proposed in this regulatory guide. The regulatory guide analysis procedures for Service Levels A and B were taken from Appendix K. However, the guide provides procedures applicable to Service Levels C and D. The regulatory guide provides specific guidance on appropriate material properties and on selection of transients for consideration, whereas Appendix K does not provide these procedures and guidance. Without this guidance, each affected licensee would have to develop appropriate procedures for Service Levels C and D, justify the choice of transients, and develop plant-specific material properties.
It is estimated that without the guidance of this regulatory guide, developing plant-specific procedutes and material Upopties and applying them to check and report the analysis results would require an additional 6 staff-months (1040 hours0.012 days <br />0.289 hours <br />0.00172 weeks <br />3.9572e-4 months <br />)
for each affected licensee. Assuming that half of the affected licensees either belong to owners! groups or could make use ofcoimon data, the total additional burden on the licensees that would be incurred by plant-specific analyses is estimated as 9 plants x 6 staff-months per plant, or 54 staff-months (9360 hours0.108 days <br />2.6 hours <br />0.0155 weeks <br />0.00356 months <br />).
In addition to the increased burden on the licensees, it is estimated that an additional 1.5 NRC staff-month would be required to review each plant-specific submittal. Thus, the total increased burden on the NRC staff assuming that half of the affected plants can be grouped, is estimated to be 9 plants x 1.5 staff-month per plant, or 13.5 staff-months (2340 hours0.0271 days <br />0.65 hours <br />0.00387 weeks <br />8.9037e-4 months <br />).
This estimate assumes that there would be only minor discussions with the licensees.
4.2 Take No Action As discussed in Section 3.2 above, taking no action is judged to be a nonviable alternative.
5. DECISION RATIONALE K
It is rehmnded tat the reglatory guide be issued because it would offer a comprehensive set of acceptance criteria, evaluation procedures, and material properties that can be used to perform the analyses required under Appendix 0 to 10 CFR Part 50 for those pressum vessels that have Charpy upper-shelf energy of any beltline material that falls below 50 ft-lb.
Issuing the regulatory guide is recommended over the alternative of endorsing Appendix K to ASME Section XI because Appendix K does not currently include (1) analysis procedures for Service Levels C and D, (2) guidance on selecting the transients for evaluation, or (3) details on tempesture-dependent material properties. Further, it is estimated that preparing plant-specific analyses that include the procedures and data that are not addressed in Appendix K would require approximately 54 staff-months of effort for the industry and approximately 9 staff-months for the NRC to review the additional information.
The NRC staff considered the possibility of worling with the ASME Code Section XI working group to modify Appendix K to include the missing procedures and data. However, given the number of plants that could need the guidance in the near term, and given the ASME codification process and the NRC's process for endorsing ASME documents, the time needed to modify and endorse Appendix K was judged to be excessive,.
The efficacy cdw procedures in the regulatory guide was demonstrated by generic bounding calculations3 performed by the NRC staff in preparing SECY-93-048. These calculations demonstrated that the requirement in Appendix 0 to
10 CFR Part 50 to demonstrate margins equivalent to those in Appendix O to Section Ill to the ASME Code could be satisfied for materials with Charpy upper-shelf energy less than 50 ft-lb for all the generic vessel geometries and material combinations considered.
3 Charles Z. Seapan. Jr., NRC, Memorandum to Jack Stunider, NRC, January 15, 1993, "Generic Bounding Analyses for Evaluation of Low Charpy UperýME .ffeon Safety Margi Against Fracture of RPV Beltine Plate and Weld Materials', Charles Z. Serpan Jr.. NRC. Memorandum to Jack Strmider, NRC, Fdxuay s, 1993, "Additional hrmation Regarding Results of Generic Bounding Analyses for Evaluation of Pressure Vessels Fabrcated Using Low ChaW Uppr-Shelf Energy Materials." Copies are available for inspection or copying for a fee firom the NRC Public Document Room at 2120 L Street NW., Washington, DQ, the PDR's mailing a is Mail Stop .L6.Washington. DC 20555; telepo (202)634-3273; fax
(202)634-3343.
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The regulatory guide acce-,.snce criteria were taken directly from the ASME efforts. The criteria were developed by the ASME Code Section XI working group over an I I-year period and represent the collective judgment of a body of cxerts representing the NRC stafl research contractors, nuclear utilities, nuclear power plant vendors, consultants, and S academia. Similarly, the evaluation procedures for Service Levels A and B were developed by this group. The procedures in the regulatory guide for Service Levels A and B are essentially identical to those in Appendix K to ASME Section 3.
Thus, the aoceptance criteria and die evaluation procedures for thc service levels that generally control the analyses are based on the consensus technical 6pinion of a large group of technical cxpes and were developed over an extended period.
The evaluation procedures for Service Levels C and D were developed by the staff and build on the procedures for Service Levels A and B. As part of a continuing effort by die ASME Section XI working group, the NRC staff has compared
1he regulatory guide procedures to odr procedures that are being developed by various organizaticns. The comparison was very favorbl; with the procedures proposed in the regulatory guide predicting lower acceptable Charpy upper-shelf energy values than would be predicted by the other procedures, which were less rigorous and, consequently, more conservative.
The procedures for transient selection are based on procedures that have already been endorsed by the star Alternatively, generic bounding transients can be used ifjustified.
The guidance on material properties is based on a state-of-the-art statistical evaluation of all available fracture touglmess data. A broad range of alternatives is offered in the regulatory guide so that methods acceptable to the staff arc offered for virtually every siuation and combination of circumstances.
The regulatory guidc provides timely, cost-effectivc guidance that is based on the consensus of a lar group of tecdincal experts represnting divers*badcgmrmds and nerecst The specific guidance is comprehensive and would provide an effective and definitive approach to performing equivalent margin anaiyscs.
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