LD-83-053, Forwards Basis for Design of Plant W/O Pipe Whip Restraints for RCS Main Loop Piping, to Demonstrate That guillotine-type Failure of RCS Main Loop Piping Need Not Be Considered in Design Basis for Sys 80
| ML20076J018 | |
| Person / Time | |
|---|---|
| Site: | 05000470 |
| Issue date: | 06/14/1983 |
| From: | Scherer A ABB COMBUSTION ENGINEERING NUCLEAR FUEL (FORMERLY |
| To: | Eisenhut D Office of Nuclear Reactor Regulation |
| References | |
| LD-83-053, LD-83-53, NUDOCS 8306200254 | |
| Download: ML20076J018 (121) | |
Text
{{#Wiki_filter:C-E Power Systems Tel 203/688-1911 Cornbustion Engineenng. Inc. Telex 99297 1000 Prospect Hill Road Windsor. Connecticut 06095 H POWERSYSTEMS Docket No.: STN-50470F June 14, 1983 LD-83-053 Mr. Darrell C. Eisenhut, Director Division of Licensing U.S. Nuclear Regulatory Commission Washington, D.C. 20555
Subject:
Basis for Design of Plant Without Pipe Whip Restraints
Dear Mr. Eisenhut:
Conbustion Engineering (C-E) has prepared the attached report, " Basis for Design of Plant Without Pipe Whip Restraints for RCS Main Loop Piping", to demonstratethatguillotinetypefailureofReactorCoolantSystem(RCS) gain loop piping need not be considered in the design basis for the System 80 NSSS. To substantiate our conclusion, the report presents: (1) a deterministic fracture mechanics analysis to validate the " leak-before-break" failure scenario, (2) details of the support system design to demonstrate that there is sufficient design margin for seismic excitation, and (3) the limits of acceptability for key parameters used in the analysis. CESSAR-F and the System 80 design have considered the double-ended guillotine break as a design basis and thus include pipe whip restraints and component stops for the RCS. Based on our new analysis, however, we believe that pipe whip restraints and component stops for the RCS are not required. While the attached report is not necessary to validate the adequacy of CESSAR-F, Staff review and acceptance of our analysis will provide a utility referencing this material significant reductions in both operating and construction costs, as well as in radiation exposure, without a concomitant reduction in plant safety. We therefore request Staff review of the attached report. C-E will formally incorporate the attachment into CESSAR-F as Appendix "C" in a future amendment . Very truly yours, COMBUSTION ENGINEERING, INC., 8306200254 830614 -_ 4 PDR ADOCK 05000 A A. E. cherer Director Nuclear Licensing AES:tmr xc: G. Meyer (Project Manager /USNRC) w/ attachment g6 i I i
Basis For Design Of Plant Without Pipe Whip Restraints For RCS Main Loop Piping Attachment to LD-83-053
. .. . _ _ - _ - _ _ _ _ _ _ _ _ __ . _ _ _ ___ ___s
Attachm:nt to LD-83-053 Pags 1 6/14/83
1.0 INTRODUCTION
In accordance with existing regulations, the design basis for CESSAR-F includes postulation of guillotines at specific locations in the Reactor Coolant System (RCS) main loop piping. In order to mitigate the consequences of these postulated pipe breaks, CESSAR-F includes pipe whip restraints and component stops which limit the blowdown area resulting from postulated guillotine breaks. Existing regulations do not address the failure mechanisms which would lead to postulated guillotines and, therefore, do not credit those features in the design of the RCS which would prevent the formation of such guil1otines. This appendix addresses those features in the RCS design which demonstrate that guillotine type failure mode need not be addressed and that pipe whip restraints and component stops for the RCS are not required. The following key issues are addressed in this appendix:
- 1. Deterministic fracture mechanics analyses to show that leaking cracks are detectable long before they may lead to pipe breaks.
- 2. The support system design includes sufficient margin to demonstrate that the characteristics of the support system remain
! valid during an SSE. That is stress resulting from combined g effects of normal operation and SSE are below the yield stress of the supports.
- 3. Key parameters which bound the applicability of this appendix are i denti fied.
In addition, probabilistic analyses relative to direct and indirect Double Ended Guillotine Breaks (DEGB) for applicants wishing to use this appendix are being performed at the requst of NRC research. 2.0 Reduction in Cost and Radiation Exposure p The total cost of engineering, materials and initial installation of pipe whip restraints for the RCS main loop piping is estimated at $3.5 million per plant. This includes: l Engineering $1.0 Million 1 Materials $1.6 Million Installation $900,000 t This cost does not include the cost associated with design of other systems associated with interferences due to pipe whip restraints. i
.I
Attichment to LD-83-053 Paga 2 6/14/83 In addition costs and radiation exposure associated with removal and reinstallation of pipe whip restraints for Inservice Inspection are estimated at $20,000 and 40 man-rems per restraint per ISI. Fourteen of the twenty-two restraints would interfere with ISI. The locations at which ISI is to be performed is plant specific; it is likely that a number of restraints will have to be removed at each ISI. Assuming that each restraint has to be removed once, the annual cost and exposure would be $7000 and 14 man-rem. For a total of $280,000 and 560 man-rem over forty years. The estimated annual cost does not include costs associated with increased outage time which may be necessitated by removal of these restraints. 3.0 Pipe Break locations Breaks in the RCS main loop piping are currently postulated at all terminal ends. Location Flow Area (in.2) Reactor Vessel Outlet Nozzle 100 Steam Generator Inlet Nozzle 600 Steam Generator Outlet Nozzle 592 RC Pump Suction Nozzle 430 RC Pump Discharge Nozzle 480 Reactor Vessel Inlet Nozzle 100 Suction Leg Slots (2) 532 r'gures 3.1-3.6 present the geometry, size and material at each pipe break location. 4.0 Support System 4.1 Support Description 4.1.1 Reactor Vessel Supports The reactor vessel is supported by four vertical columns located under the vessel inlet nozzles. The pad at each nozzle provides a surface i to which the colurn is bolted. This pad also acts as a horizontal key l to positively locate the vertical centerline of the vessel. It mates with a welded plate structure embedded in the concrete and allows free radial growth of the vessel during thermal expansion while supporting l the vessel horizontally during earthquakes anc following a postulated j pipe rupture. l The vertical columns are designed to support the vessel and resist vertical motion during earthquakes and following postulated pipe rupture. The columns are manufactured of SA 508 material. This material exhibits characteristics which minimize radiation embrittlement.
Attachment to LD-83-053 Pagt 3 6/14/83 At the bottom of each column is a baseplate which is drilled to accept anchor bolts. Shear bars integral with this plate and preloaded anchor bolts are the mechanism by which coluen loads are transmitted to the foundation. The baseplate also acts as a keyway for a horizontal key welded to the lower vessel head. An energy absorbing material is used between the key and keyway to provide horizontal seismic support, while limiting the load on the vessel head during a postulated pipe rupture. The bolts at the top and bottom colunn flanges are preloaded to insure no separation during the maximum postulated loading conditions. Thus the system stiffness (and corresponding natural frequencies), is derived from the column stiffness and not that of the bolts. This system was designed primarily for abnormal conditions. An important feature is that the vertical load is transferred to the basemat directly rather than through the cavity wall which is already subjected to high horizonal loads. Note that the low friction bearing can be shimmed after the structure has been installed, thus allowing recovery from errors and also convenient adjustments. 4.1.2 Steam Generator Supports The steam generator is supported by a conical skirt welded to the steam generator lower head. The skirt provides a bolting surface for a heavy steel sliding base. The sliding base rests on four plates. During heat-up, the reactor vessel supports expand more than the steam generator supports and thus the front of the generator is liftec and the weight shifts to the rearmost support. Machined cutouts in the base act as keyways for embeded keys which support the generator horizontally. The base slides on four low friction spherical head bearings; low friction bearings are also used in the keyways to minimize resistance to thermal expansion. Slotted holes are provided in the sliding base to accept anchor bolts. The clearance between the top of the sliding base and the anchor bolt nuts is set during hot functional testing to minimize the gap without restricting thermal expansion. The side bearings and shim assembly are designed to be attached to the sliding base following welding of the RCS piping and the steam piping. Thus, any rotation of the steam generator which results from welding can be accommodated by shimming. The front key also acts as a stop for the generator to accommodate postulated pipe ruptures in the hot leg. Shimming of this stop is accomplished following hot functional testing; the method of attachment of the shim allows this to be accomplished. l I Horizontal support at the top of the. generator is provided by two keys and two hydraulic snubber assemblies. They act as horizontal supports for the steem generator during earthquakes and following postulated pipe rupture, while allowing motion parallel to the hot leg due to thermal expansion. At each keyway installation, the shims which control the gap are installed after welding of primary and secondary pipes to easily accommodate any generator movement during this process. i
Attrchment to LD-83-053 Paga 4 6/14/83 Each snubber assembly consists of a lug welded to the steam generator, a lever, two links, one snubber, one clevis pinned to the lever, and one clevis pinned to the snubber. Each clevis is bolted to a built-up member or stub column which is anchored to the secondary shield wall and the bolts are preloaded. As with the pinned joints at the pump supports, spherical bearings are used at the connections of the links. This allows for transient expansion and contraction of the steam generator shell. The hydraulic snubbers allow the steam generator to move freely during gradual thermal expansion but will resist any rapid movement caused by either seismic or postulated pipe rupture loading. 4.1.3 Reactor Coolant Pump Supports The supports and stop of the reactor coolant pump control movement of the reactor coolant pump in the horizontal and vertical planes during earthquakes and following accidents, but accomodate thermal growth during plant heat up. The reactor coolant pump and motor assembly is supported by four vertical columns pinned to the pump mounting ring. It is supported for seismic and loss of coolant accident loads by two horizontal columns pinned to the top of the motor mount, two horizontal columns pinned to the pump mounting rings, and a horizontal snubber system attached to the top of the motor mount. In addition, horizontal restraint is provided, following a postulated pipe rupture, by a stop located at the pump casing at the elevation of the discharge nozzle. Each column, horizcntal and vertical, and the snubber assembly end in a clevis, through which loads are transmitted to the building st ructu re. Eacn clevis is drilled to accept anchor bolts by which it is connected to an embedment. All bolts are preloaded to prevent joint separation. The pump stop mates directly with the pump casing. The clearance between the mating surface and the stop is shimmed to a specified size and verified during hot functional testing. The thermal expansion of the reactor coolant pump is not always in the same direction. To compensate for this a spherical bearing is employed at each pinned joint. Without a spherical bearing, a significant load would be imposed on each pin due to thermal movements not directly in line with the axis of the pin. 4.2 Mcterial, Fabrication, Installation and Design Margin Tables 4.2.1 through 4.2.3 present information relative to the material and fabrication of each member of the RCS support system. The material and fabrication details in the tables are generic and substantially apply to all System 80 plants although some differences may exist in the details from plant to plant. The support design includes margin for seismic excitation. In all cases the calculated stress for normal operation plus seismic is less than 70% of the allowable stress for a faulted condition.
Attechment to LD-83-053 Paga 5 6/14/83 With respect to insta'.lation, especially potential installation errors, particular care has been taken in the transmittal of design information between the affected parties in the design and construction of these supports. In addition, the installation
. receives careful review in the form of gap verification during the installation, at the time of hydrostatic testing, hot functional testing and post core hot functional testing. Since anchor bolts are preloaded, they provide verification of adequacy of both the bolt and the embedment to which they are attached.
5.0 Fracture Mechanics Analyses The enclosed report describes the fracture mechanics analyses in support of leak before break. The report contains descriptions.of analyses with nominal and significantly degradec material properties and shows that significant margin between a detectable crack and one which could result in a pipe break exists. 6.0 LEAKAGE DETECTION Reactor Coolant Pressure Boundary Leakage Detection System Requirements are defined in Section 5.2.5 of CESSAR F. 7.0 Key Parameters The key parameters of this appendix which define the limits of acceptability of the analysis presented in this appendix are: Nozzle Materials - SA 508 CL 1, 2, or 3 Safe End Materials - SA 508 CL 1, 2, or 3 Piping Material - SA 516 Gr70 Operating Pressure - 2250 Psi Piping Dimensions: 42" ID (Hot Leg) 3 3/4" (Pipe) 41/8" (Elbow) 30" ID (Cold Leg) 2 1/2" (Pipe) 3" (Elbow) Support Margin - Fauled Allowable 1 1.5 N.O. + SSE Leak Detection Capability Section 5.2.5 of CESSAR F. Seismic response spectrum at RC Pump and Reactor Vessel enveloped by spectrum in Figure 41 of this report. F45404
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Attachment to LD-83-053 Page 7 6/14/83 SAFE END I l V f PIPE . PIPE 10 42" THICK?iESS = 41/8" (Elbow) N0ZZLE MATERIAL - SA 508, CL1, 2, or 3 PIPE MATERIAL - SA 516, Gr. 70 ' SAFE END MATERIAL - SA508 CL1, 2, or 3 STEAM GENERATOR IllLET N0ZZLE FIGURE 3.2
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Attechment to LD-83-053 Page 19 6/14/83 i Leak Before Break Evaluation of I the Main Loop Piping of a CE Reactor Coolant System j Combustion Engineering, Inc. - 1000 Prospect Hill Road Windsor, CT 06029 f i I i I i h I [ l i
Attcchment to LD-83-053 Page 20 6/14/83 Table of Contents
- 1. ABSTRACT
- 1. INTRODUCTION
- 2. LEAK BEFORE BREAK CRITERION
- 3. THROUGH WALL CRACK GROWTH
- a. Initial Crack Size and Loadings
- b. Crack Growth Analysis
- 4. CRACK LEAKAGE
- a. Crack Opening Area Analysis
- b. Leak Detection
- 5. CRACK STABILITY ANALYSIS
- a. Crack Stability Criteria
- b. Normal Operating Loads
- c. Seismic Loads
- d. Axial Slot in Elbow
- 6. MARGINS ON-System 80 MAIN LOOP PIPING
- a. Margin on Seismic Load Capability
- b. Margin on Unstable Crack Size
- c. Margin on Axial Slot In Elbow
- 7. CONCLUSION
- 8. REFERENCES
- , - . ~ > - - - . - ,n., - , , .-e
Attachment to LD-83-053 88e 21 6/14/83 ABSTRACT The studies conducted by CE over the past decade ralated to the demonstration of leak-before-break in the main loop piping of a CE PWR are summarized. Recent analyses which address the safety margins available in meeting the leak-before-break criterion are also presented. The results clearly demonstrate that the margin against instability for detectable leaking cracks is i substantial in terms of margin on material properties, loadings, and crack size required for instability for both normal operation and seismic conditions. All the requirements' necessary for the demonstration that a leak-before-break condition exists in the main loop piping of a CE PWR are shown to be satisified with considerable margin. i 1 I l
- Attcchment to LD-83-053
- 1. INTRODUCTION Nuclear power plants are designed to withstand very large mechanical loadings that are intended to envelop conservatively a wide range of hypothetical accident conditions. These loadings are typically associated with a hypothetical initiating event that is judged to be more severe than any realistic event. The primary coolant system main loop pipe break is one of these initiating events that forms the design basis for many systems
!. and components of a pressurized water reactor (PWR). The present criteria; which define the location; type and size of pipe breaks are based on the assumption that a sudden complete circumferential severance (guillotine break) of a pipe can occur Ref. -(1). Advances in the field of fracture mechanics and elastic-plastic stress analysis during the
./
_past decade have provided the capability to more realistically assess the way that cracks would grow in piping systems. Application of these analytical methods to the main loop piping in the CE PWR has demonstrated that a leak before break condition exists and that a complete circumferential severance of a main loop pipe cannot occur.. 1 The analyses that have been performed in support of the leak-before-break demonstration 'are described in this report. These analyses form the basis ! for the position that the existing criteria could be revised to not require the consideration of the guillotine break, .and still satisfy tne requirement of enveloping conservatively all possible accident conditions. Analyses are also presented which address the leakt before break condition of axial slots in piping elbows. These analyses demonstrate that a large margin exists for leak-before-break.
- This analysis emphasizes circumferential pipe cracks because the
- circumferential4 seyerance currently hypothesized is the most severe type of
- pipe break and~results in the requirement for pipe whip restraints.
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-Many CE. published reports' are _ reviewed and summarized in order to bring all
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importaat CE work on the leak-before-break issue together in one report. Frequently,,se::tions of these reports are restated " verbatim" with permission of t!te authors of the original' studies. -
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Attachment to LD-83-053 Page 23 6/14/83
- 2. LEAK-BEFORE-BREAK CRITERION In order to demonstrate that a leak-before-break condition exists for a particular piping system, three conditions must be met. The first condition is that any initial flaws must tend to propagate through the pipe wall rather than in the circumferential or axial direction. This characteristic of the flaw growth is dependent on the loading and environmental conditions to which the piping is exposed. The second condition is that through wall cracks must open sufficiently to allow detection by normal leakage monitoring under nornal full power loading conditions. The nature of the crack opening is dependent on the piping stiffness, the normal operation loadings and the material properties of the pipe.
The third condition is that cracks of detectable length must remain stable even under severe loading. The most severe loading is considered here to be the Safe-Shutdown Earthquake or SSE seismic loading. Crack stability is dependent on the toughness or crack resistance of the material, and the manner in which the cracked pipe is able to distribute or shed loads as y crack extension occurs. a Leak-before-break can be demonstrated if the three conditions identified above are satisfied with sufficient margin to assure that unforeseen variations in loadings or material properties cannot cause one of the conditions to not be satisfied. Analyses that have been performed by CE over the past decade addressing these three conditions are summarized in sections 3, 4, and 5. In section 6, more recent analyses aimed toward quantifying the safety margins in meeting the three conditions. 1 { L----_. . . . . ,
Attichment to LD-83-053 Page 24 6/14/83
- 3. THROUGH WALL CRACK GROWTH The first step in the leak-before-break evaluation is the determination of the manner in which a crack could grow through the pipe wall and cause a violation of pressure boundary integrity. Crack growth is caused by cyclic loading on an existing crack. In order to evaluate crack growth, the loading conditions and the initial crack size must be established.
- a. Crack Size and Loadings The anticipated loading conditions for CE PWR primary piping are shown in Table 1. These loading conditions are well understood and contain no severe thermal or dynamic conditions. These anticipated conditions are employed in the ASME code fatigue analysis for the piping. There is the possibility that a more severe loading, e.g. thermal transient, may occur during an emergency or faulted type of event, but the scarcity of such events precludes the need for consideration of such loadings in a fatigue or fatigue crack growth analysis.
The main loop piping has no valves which might open or shut, it experiences no rapid thermal transients during normal operation and it is not subject to significant flow stratification during normal operation. Clearly, since these statements cannot be made about piping in general, the conclusions drawn concerning leak before break apply only to the main loop piping. The CE main loop pipes are fabricated and inspected in accordance with NB 2532 and NB 5000 of the ASME Section III Code which requires volumetric examination of 100". of the base metal and weld joints and allows indications no longer than three inches nor deeper than 10". of the pipe wall thickness. Therefore, only small cracks could exist in the piping before service. In order to conservatively evaluate crack growth and extension, a variety of crack sizes much larger than those expected to exist are considered in the subsequent analyses.
- b. Crack Growth Analysis Crack Growth analyses have been performed based on the methods of linear elastic fracture mechanics, Ref. (3). Recent analyses have employed the methods of the ASME Code Section XI as those f racture mechanics procedures have developed Ref. (4).
I i Section XI of the ASME Code, Ref. (5) defines a fatigue crack growth rate law of the form: da dTi
=
C'(dK)" g (1) where n is the slope of the log (da/dN) vs. log ( K y) curve, and C' is a scaling constant. This material property curve has been determined experimentally, and the material constants for carbon steel l C' = for fatiguegrack andgrowth n = in a water 3.726. The environment rate of crack are as (da/dN) growth follows: is 3.795 x 10~ measured in inches per cycle from this relationship. This crack growth law is intended to be a conservative upper bound to the a
Atttchment to LD-83-053 Pegn 25 6/14/83 experimental data, however, recent fatigue crack growth studies have produced data which lie above this curve. Figure i shows the da/dN vs. Kg curve which has been proposed as a revision to the Section XI curve and is seen to envelope all of the fatigue crack growth data. The results of this study include the upper bound to the crack growth rate curve as given in Figure 1. For the purpose of determining the range of defect sizes which could grow to threaten the integrity of the system, semi-elliptical shaped inner surface flaws were hypothesized for various initial crack depths, ao, and lengths, 2 C .oA computer methodology was used to evaluate the stress intensity factor for a given flaw size and loading function and then compute the growth rate in both the through thickness and circumferential directions of the flaw under cyclic loading conditions. The method of analysis is based on the Section XI, Appendix A flaw evaluation procedure. This method is applicable to flaws which have not fully penetrated the wall thickness. From this method for Kg determination, the O Kg level is calculated based on the crack size and loading conditions. Using a stepping procedure for the number of cycles of loading in a given time period, depth and length crack growth rates are calculated and the corresponding change in crack size is determined. This permits a determination of the time to produce first leak when an existing flaw would enlarge and subsequently " pop-through" the thickness of the pipe. This " pop-through" phenomenon is what is meant by a suddenly-appearing througn-thickness crack. The circumferential length of the through-wall crack is important to the determination of crack stability. For cracks which are calculated to grow to penetrate the wall thickness, the subsequent calculation of the stress intensity factor, Ky, was evaluated using the finite element method since the Section XI flaw evaluation procedure does not extend to through-wall cracks. The finite element analysis enabled the determination of Ky as a function of the circumferential crack length, 2C, and the applied load. This information was incorporated into the crack growth procedure. In this manner the fatigue crack growth study was continued for circumferential growth of through-wall cracks. A wide range of initial flaw sizes and shapes were considered, and the resulting crack growth rates were calculated for the design basis loading transients and corresponding frequencies of occurrence given in Table 1. An example of the predicted growth behavior for a defect with an initial depth of 0.5 in. and an initial lengtn of 39.0 in. is shown in Figure 2. Under the influence of tne prescribed cyclic loading history such a flaw was calculated to become a leaking crack in 21 years of operation, which is less than the normal (40 year) plant life. The circumferential extension of the crack is only about one inch which is negligible compared to the initial crack length. It should be emphasized that the loading histories used in this analysis are conservative representations of the design transients and are intended to describe the upper limit of possible < eactor operating
Attachment to LD-83-053 Paga 26 6/14/83 experience. If, after formation of a through-wall crack, the defect remains undetected and operation of the reactor continues, the calculated f atigue crack growth is also shown in this figure. It can be seen that the through wall crack would remain stable for many more years of service, increasing the likelihood of detection. Figure 3 shows a similar plot of crack size vs. years of operation for an initial flaw with dimensions a o = 1.0 in and 2 Co = 34.0 in. This initial flaw size also results in the formation of a large circumferential through-wall leaking crack in only 4 years of reactor operation. The circumferential extension of this crack is negligible. Subsequent extension of the leaking crack due to continued operation is also shown on Figure 3. The results of the fatigue crack growth study for a .35 in. deep and 45.5 long initial flaw are presented in Figure 4. From this figure it is seen that the time required to cause a leaking crack would be 38 years. Again circumferential extension even of such a long (half circumference) crack is negligible. In earlier CE work', Ref. (3), it was shown that cracks shorter than those considered above would require many more than the design basis number of cycles of loading to grow through the pipe wall. The startup-shutdown transient was found to be the greatest contributor to the usage factor ior the main loop piping. A cyclic stress, conservatively enveloping this startup-shutdown stress was applied to hypothetical flaws one inch deep and from 8 to 18 inches long in the circumferential direction in both the 42-inch diameter hot leg and 30-inch diameter cold leg piping. Figure 5 shows that the number of start-up-shutdown cycles necessary to cause a one-inch deep crack to grow through the pipe wall and leak is an orcer of magnitude greater than the plant life. In these cases it was observed that the circumferential crack extension was small. The analyses indicate that the transients that the pipe is anticipated to experience produce preferential crack growth in the through-wall direction of the pipe thereby causing a leak prior to significant circumferential extension. In Reference 3 the same conclusions about preferential through thickness crack growth were also demonstrated for cracks in the axial direction. One other crack extension mechanism, stress corrosion cracking has the potential to cause crack growth more uniformly around the pipe ci rcumference. The CE main loop piping environment, however, is not corrosive and no evidence of stress corrosion in this piping has ever been observed. This crack growth mechanism is not considered to be active in CE main loop piping.
Attachment to LD-83-053 Pagn 27 6/14/83
- 4. CRACK 1.EAKAGE The second step in the leak-before-break evaluation is the determination of the amount of leakage which will result from a given crack which has extended through the pipe wall . CE has performed a detailed study of crack opening areas- at two locations in the main loop piping: the pump discharge leg terminal end at the reactor vessel inlet nozzle.and the hot leg terminal end at the reactor vessel outlet nozzle, Ref. (6). These locations are selected because they are regions of relatively high stress and are locations where guillotine ruptures must be postulated according to the present pipe break criteria, Ref. (1).
The crack opening area was calculated for normal operating conditions since these are the most prevalent conditions during plant operating life, when leakage is to be detected. Since the leak rate is related to the amount of crack opening, a finite element analysis was performed to calculate crack opening areas as a function of crack length for various orientations around the pipe at these locations.
- a. Crack Opening Area Analysis The crack opening area was determined for several circumferential through-crack lengths oriented at the top, bottom, and sices of the discharge leg and hot leg terminal ends at the reactor vessel nozzles under normal operating conditions. Figure 6 shows the locations in the primary coolant system of the regions analyzed. By symmetry of geometry, material, and normal operating loading, the results for these two regions also apply to the other pipe terminal ends at the reactor vessel nozzles. Figures 7 and 8 show the geometry details used for the finite element modeling of the discharge leg and hot leg, respectively, and the coordinate systems in which the forces and moments are specified. Table 2 gives the combined forces and moments due to pressure, weight, and thermal expansion for steady state normal operating conditions. All crack opening area analyses were linear elastic using the material property values which are given in Table
- 2. Plasticity is conservatively ignored since plastic deformation would cause greater opening areas and more leakage. The MARC finite element program was used to perform all crack opening area analyses.
The extent of the detailed model of both structures was chosen so that nozzle and pipe / elbow end modeling would have a negligible ef fect on crack opening area. Figure 9 shows the finite element mesh for the 30 inch discharge leg terminal end with a crack.' The 42 inch pipe mesh is not shcwn because it is very similar. Figure 10 shows the mesh refinement in the structura imediately surrounding the crack which was needed for accurate crack opening area determination and simulation of cracks of different length without the need for overall mesh regeneration. For each terminal end finite element model, the total combined "in-system" loads of Table 2 were applied to produce equivalent beam displacements and rotations at the ends of the models. These calculated displacements and rotations were then applied to the crack
Attachment to LD-83-053 Pagn 28 6/14/83 models because the displacement-controlled loading was judged to realistically represent the nature of the actual normal operating loading applied to the cracked structures. The boundary deformations of these structures, even under loadings like axial pressure, are substantially constrained by the resisting stiffness of the rest of the piping-component system. For determining leakage rates from a narrow through-thickness crack in the pipe, it is necessary to determine the minimum opening area at any location through the wall because that section limits the flow from such a crack. For this reason, the crack opening area was calculated at the outside surface, inside surface, and midplane of the pipe wall. The midplane crack opening area was calculated directly from the shell midplane displacements. For the inside and outside surface areas the relative displacements due to the through thickness rotations were added to the midplane displacements. The opening area computed at the midplane of the pipe wall was compared to the crack opening area at the surfaces. For the hot leg the smallest area was always greater than 92 percent of the midplane value for all crack sizes and orientations evaluateo. For the discharge leg the smallest area was always above 73 percent of the midplane value for all crack sizes at all orientations, except the bottom. At the bottom, the area was as little as 23 percent of the midplanc value for the shortest cracks. These results illustrate that there are significant variations in the crack opening area at the inner and outer surfaces of the pipe, especially for the minimua crack opening area orientation. Figure 11 shows the minimum pipe surf ace crack opening areas vs crack length at the various orientations around the discharge leg terminal end. The bending moment at the terminal end produces a greater opening area on the top of the pipe than on the bottom during normal operating conditions. Cracks at all pipe orientations open significantly indicating that the axial load, caused mostly by system pressure, predominates over the bending moments. Figure 12 shows the minimum pipe surface crack opening areas vs crack length at the top and side of the hot leg terminal end. Cracks hypothesized in the buttom of this pipe do not grow or open because the region is in compression.
- b. Leak Detection There are two major facets to crack detection based on leakage in addition to the crack opening size. These are the leak detection sensitivity, and the flow rate correlation for leakage througn a crack. A more detailed discussion of leakage rates is presented in Ref. (7).
For a PWR in the USA leak detection systems capable of detecting less than 1.0 gallon per minute (gpm) leakage from the primary system, with a Technical Specification upper limit of 10 gpm are employed per the guidance of Regulatory Guide 1.45 Ref. (8). Diverse measurement means are utilized, including water inventory monitoring, sump level and flow monitoring, and measurement of airborn radioactive particulates or gases.
Attachment to LD-83-053 Pags 29 6/14/83 The other major facet of crack detection based on leak rate, namely the flow rate correlation for leakage through a given crack size, can not be predicted precisely. Variables such as surface roughness of the side walls of the crack, the nonparallel relationship of the side
. walls due to the elongated crack shace, and possibly zig-zag tearing of the material during crack formation all introduce uncertainties in defining an exact flow rate correlation.
NUREG/CR-1319 Ref. (9)', provides a treatment of leakage through small cracks considering various uncertainties in crack definition, including crack wall surface roughness, effective 1./Dh ratio of the elongated crack shape, and the possibility that the crack may be longer at the inside of the pipe wall than at the outer surface of the pipe, resulting in a convergent opening. The results of Figure 4-13 of Reference (9) for typical PWR conditions at 2250 psi and 550*F for a high friction factor of .01 are replotted on Figure 13 in units of gpm per square inch of crack opening versus outer surface crack area, Ae. Also plotted in Figure 13 are flow predictions based on similar orifice flow with a discharge ccefficient of 0.6, and also a flow prediction using a Henry-Fauske, Ref. (10), critical flow model. The Henry-Fauske correlation was developed on the basis of subcooled flow through nozzles, and provides an upper bound for flow through an irregular crack opening. The orifice flow does not consider , subcooled water effects, and the constant discharge coefficient does not consider the irregular crack shape. Even so, the orifice prediction falls in the range of the NUREG/CR-1319 predictions, providing a measure of comparison. The NUREG/CR-1319 predictions show a slight increase in flow rate per unit of exit plane area with increasing area, ano o large increase for decreasing Ae/Ao ratios. Since, for the purposes of identifying a through wall crack by means of leakage it would be conservative to underpredict the flow rate, the lowest value of all of these various predictions is recommended. The lowest flow rate prediction is about 885 gpm/sq. in, at .001 sq. in. This means that a crack which opens to slightly greater than .001 sq. in. will leak at least 1 gpm, and, therefore, will be within the range of detection by the normal leakage monioring systems for a PWR. Using the relationship between leakage rate and crack area of 885 gpm per square inch, the leakage rate for all the crack cases of Figures 11 and 12 can be determined. These flow rates are shown in Figure 14 Considering some margin for conservatism, through wall cracks resulting in leakage of 1 to 2 gpm will be detected during operation. This limit is also shown on Figure 14. It can be seen that relatively small cracks produce detectable leakage. For all locations except the bottom of both terminal ends, a 5-inch long crack would be detectable. For the bottom of the discharge leg terminal end a 10 inch (254 mm) long crack would be detectable. The crack length required for detectable leakage at the bottom of the pipe is greater than the other orientations because the normal operating loading is smaller in this region. This smaller loading is seen later when the stability of cracks of various orientations is considered. _ l
Attechment to LD-83-053 Prg2 30 6/14/83 The bottom of the hot leg terminal end is always in compression and is not considered a viable crack location. . . . . . . m
Attachment to LD-83-053 Pagn 31 6/14/83
- 5. Crack Stability Analyses The third condition for leak before break is that cracks which are large enough to leak will remain stable during extreme loading conditions, thereby preventing a complete pipe severance. Sufficient margin on stability must be available to account for variations in actual leak detectability, material properties and loading events between the time of the beginning of leakage and plant shutdown for leak repair.
! CE has performed fracture mechanics analysis on a variety of hypothetical crack sizes subject to a variety of loading conditions. Both normal operation loads and seismic loadings have been considered to determine the size of crack which will remain stable under various loading conditions.
- a. Crack Stability Criteria Two crack stability criteria have been used over the years to assess the likelihood of unstable crack extension in the main loop piping.
These methods are the traditional linear elastic fracture mechanics which employs a KIC criterion, and the Ductile Tearing or J-integral ductile fracture mechanics which employ a JIC and Tearing Modulus criteri on. In linear elastic fracture mechanics, the stress intensity f actor at the tip of a hypothetical crack, K ,g is computed as a function of loading and geometry. This factor represents essentially the crack ! opening force applied. This value is compared to the fracture toughness, KIC which is the material resistance to fracture. If Kg < K IC then the crack is stable and if Kg 2, KIC, then unstable crack extension occurs. The ductile fracture mechanics methods have been developed more recently and, therefore have been used in various stages of that development in the analyses described here. The tearing modulus concept is an elastic plastic crack instability theory based on the J integral elastic plastic crack tip parameter and a J-resistance curve material property such as shown in Figure 15 Ref. (11). This figure shows the results of a series of tests indicating the amount of crack extension as a function of the value of J at the crack tip. The J value below which there is essentially no crack extension is called i J IC* The slope of the line beyond JIC giving the rate of increase of J required for subsequent crack extension, 3J/ ba, is used to assess the stability of the crack. Figure 16 is an idealization of the J-resistance curve which illustrates the instability criterion. If the loading on the crack is such that the rate of change of J with crack extension}J/ b a applied is greater than the resistance &J/ ha material, then unstable crack extension will occur.
Attcchment to LD-83-053 Pega 32 6/14/83 If hd/h a applied is less than )J/cl a material then unstable crack extension will not occur even though J is exceeded. The non-dimensional tearing modulus T, is defihd as: T = )J E 2 (2) d Co where E is Young's modulus and Io is the yield stress of the material . Figure 17 shows how the point of instability can be found as the intersection of the loading curve in terms of J(T) and the J resistance material property curve. This figure illustrates how the structural behavior influences crack instability. If the J applied does not increase with crack extension more than the J-Resistance curve because of load redistribution or load shedding, i.e. T a is small, then the crack will not become unstable. Ifhowever$ plied the crack loading' J applied increases rapidly with crack extension i.e. T applied is large, instability will occur. Confirmation of the applicability of the JIC curve of Figure 15 to the piping material used in actual plants can be attempted by comparison with actual material data. The material data obtained for each pipe section, however, is typically limited to Charpy tests so a Charpy/KIC/dIC correlation must be employed. Figure 18 shows the Charpy energy vs temperature data for Palo Verde Unit 1 SA516 Gr70 pipe material (in plate form). Figure 19 shows the Charpy energy for weld material for a SA516 Gr. 70 to SA533 B1 or 508 Class 2 weld, typical of the pipe to component safe end welds. Using the Barsom-Rolfe-Novak correlation X IC = 2E (CV N ) / (3) where CV N is the Charpy energy, and the plane strain relationship JIC
- KIC II V 2I (4)
E i whereV is Poisson's ratio, the resulting XIC and JIC vs temperature relationships are shown in Figure 20 and 21 for the base metal and weld material. I Figure 21 indicates that the J r value of 600 inib/in.2 (of Figure
- 15) is reasonably conservative har the actual pipe material at operating temperatures. The figure also indicates that the weld material has a significantly higher JIC and use of the pipe material J
IC for stability evaluations will be very conservative for welds. For the cases of linear elasticity and small scale yield fracture mechanics the parameter J is related to Kg by Kg = d d . E/(1-y 2)' for plane strain (5) K =VJ.E ' for plane stress
Attachment to LD-83-053 Pags 33 6/14/83 This relationship is frequently used with the finite element method for the calculation of Ky . It also permits a comparison of the two crack stability criteria. For example, from Figure 15, JIC is found to be about 600 in-lb/in. for SA 516 Gr. 7g. According to equation (3), assuming plane stress, and E = 30 x 10 psi, KIC = 134 ksiffiP. It is generally accepted that linear elastic fracture mechanics
. applies until XIC approaches the upper shelf toughness which is in the 200 ksidi~n'n to 250 ksi G range. Using 225 ksi 6iPas~ an average upper ghelf toughness, the corresponding J would be 1690 in ib/in. . Figure 15 indicates that significant crack extension would result from application of a J value of this magnitude. The rate of increase of J, however, is required to assess crack instability.
- b. Normal Operating Loads Both short through wall and long through wall circumferential cracks have been analyzed for crack stability. Static analyses have been employed for small cracks to determine the margin against crack instability for cracks which may be just leaking during normal c operation Ref. (6).
The stability of through-wall cracks in primary system piping is evaluated using the J-integral technique. The crack tip parameter, J, can be compared to the experimentally cetermined elastic-plastic toughness , JIC, to evaluate the stability of a crack. For J < JIC' no crack extension will occur, hence, the crack will remain stabTe. The J-integral was evaluated in the finite element analysis using Park's method Ref. (12) which has been demonstrated to produce accurate results without the need for special crack tip elements. The J-integral value was calculated at both crack tips for normal operation loadings which include pressure, weight and thermal expansion forces, at various orientations around the circumference of the terminal end models discussed in Section 4a. From the calculated 3 l stress values, no significant plastic deformation would be expected at the crack tip under normal operating loads for crack lengths less than 25 inches or so. A plot of J vs crack length is shown in Figure 22 for through-wall cracks centered at the top and side of the hot leg terminal end and the top, outward side and bottom of the discharge leg terminal end. For all orientations, the calculated J-integral value increases as crack length increases. The J value for the hot leg It follows the pattern gssociated with a crack dominant bending load. reaches 224 in-lb/in for 22 inch at the top of the pipe. i For the discharge leg the J values are much less than the J values for j the top of the hot leg. The J value for a 171/2 incn crack at the top of the discharge leg is only 26 in-lb/in.2 The crack in the bottom of the discharge leg terminal end, which was seen to have the largest crack length for detectable leakage from Figure 14 has, by l far, the lowest J value. This indicates that the bottom is not a critical region for crack stability or concern for violation of the leak-before-break criteria.
i Attachment to LD-83-053 Paga 34 6/14/83 The computed J values for all cracks are well below the critical value of J IC in Figure 15, thereby assuring that no crack extension due to normal operating loads will occur for these cracks. A dynamic elastic plastic analysis was performed to evaluate the stability and opening area of long hypothetical circumferential cracks, in the 30-in. ID, cold leg pipe in Ref. (13). The pipe is assumed to be under normal operating pressure of 2250 psi and subject to the axial load caused by that pressure. A circumferential crack is assumed to suddenly appear with a length of half the circumference of the pipe. A schematic view of the pipe containing the circumferential crack is shown in Figure 23. Since two planes of symmetry exist, one quarter of the pipe can be modeled. The finite element model of one-quarter of the pipe is shown flattened out and not to scale in Figure
- 24. The boundary axially remote from the crack is permitted to move axially and rotate as a plane. The force on the boundary is the axial force caused by the pressure and no depressurization due to the crack opening is assumed.
The pipe material, SA516 Gr 70, was permitted to deform plastically in accordance with the stress strain curve of Figure 25. The crack The maximum opening area durigg the opening is shown in Figure 26. opening of 5.3 in. occurs at 3.0 milliseconds. This opening would result in over 5000 gpm leakage on the basis of the leak rate relationship discussed in Section 4.b. The J-integral computed during the opening is shown in Figure 27. These values were computed by MARC and are the average of two ngar crack contours. The maximum J-integral value is 1250 in-lb/in. which is above the JIC value of Figure 15. It is expected, therefore, that some crack extension would occur to elongate the crack by less than 0.1 inches. The corresponding Kg according to equation (3) is about 190 ksi VIiPwhich is below the upper shelf toughness, giving one indication that instability will not occur. For a tearing modulus evaluation the rate of change of J with crack length is requirgd. From Figure 15 A J/4 a material is about 10,000 in Ib/in, which leads to a tearing modulus at operating temperature to be: T,,e = 10,000,x 30 x 10 6 , = 480 25 x 10" x 25 x 10" The J applied of 1250 in Ib/in.2 corresponds .to a half circumference crack which has a half length, a, of 24 inches. Recent work (see Section 6.b) has indicated that Ja for this geometry can be approximated as a cubic function ofpliedcrack length, ie, and, therefore, SJ/J a applied = 3Caa 2 (7)
Attachment to LD-83-053 Prgs 35 6/14/83 Substituting equation 6 into equation 7 leads to: clJ/ da applied = 3J/a = 156 in Ib/in.2/in, and, therefore, T applied = 7.5. Since Tapplied < Tmat this tearing modulus evaluation also It is concluded, indicates that cracx instability will not occur. therefore, that the circumferential crack halfway around the pipe will not be unstable if it suddenly appears during normal operating conditions,
- c. Analysis of Crack Subject to Seismic Loads In order to determine the largest crack which would remain stable under seismic loadings, a dynamic elastic plastic finite element analysis of a crack in the discharge leg terminal end was performed (Ref.14).
In the previous section it was demonstrated that a crack must exist more than halfway around the circumference of the pipe before instability (rapid crack growth) can occur due to normal operating loads. As a first consideration for the seismic analysis, the same crack size will be evaluated for stability. Therefore, a one-half circumference stable crack is postulated to occur in the discharge leg terminal end, a region of particular concern to the integrity of the primary cooling sy stem. The crack is presumed to exist at a point in the Safe Shutdown Earthquake (SSE) loading transient wnich would produce the most severe loading condition at the crack tip. A detailed elastic-plastic dynamic analysis was performed to determine the overall response of the pipe. The stress intensity factor, K was computed as a function of time using the J-Integral technique.1,The calculated values for Kg are compared with experimental material toughness data to determine the inherent resistance of the material to further crack propagation. The maximum crack opening area is calculated in order to determine leak rates. All computations were performed using the MARC general purpose nonlinear finite element program. c.1 Main Loco Piping System Model For the seismic analysis a model representing the entire primary system is employed. A three dimensional shell model of the eloow section of the discharge leg pipe was constructed using doubly-curved thin shell elements to provide a detailed finite element description of this region. The shell element was chosen to model the pipe elbow and nozzle because through-thickness cracks can be modelled with relative ease, localized plasticity effects can be included, and the J-Integral technique can be utilized to calculate Kt at the crack tip.
Attachment to LD-83-053 ' Pagn 36 6/14/83 Because the discharge leg is a single component in a more complex system, it is important to analyze the response of the pipe to seismic
-loads considering the Interrelated effects due to the adjacent structural members. For this purpose, three-dimensional beam elements were used to model the reactor vessel and its vertical support columns, as well as the reactor coolant pump, its horizontal and vertical supports', and the snubber. Beam elements were also used for the remaining portions of the discharge leg pipe which were not modelled as shells.. A superimposed view of the finite element model of the pipe and the other system structural components is presented in Figure 28.
Only one leg of the reactor coolant system was evaluated in this analysis. The criteria for analysis of uncoupled subsystems are discussed in Reference 15. The major components and their support structures are modelled so that the substructure model-will respond as if it were part of the entire system. Seismic loading is applied as time history motion of the supports. Boundary conditions were applied to the model at the points of attachment to the foundation. The behavior of the model was checked under static conditions, with and without the presence of a crack, to verify the overall response of the finite element model to externally applied loading. A static analysis of the system containing a hypothetical one-half circumference crack was performed to determine the effects of internal operating pressure on the cracked structure. The crack was presumed to exist at the cutside of the elbow halfway around the circumference of the pipe where crack opening effects are expected to be at a maximum. To be conservative, it was assumed that the presence of the crack does not produce a depressurization of the system wnich would tend to reduce the level of stress in the pipe. The maximum static value of K at the crack tip was calculated to be 92 ksi NTT&7 This value of K due to a one-half circumferential i crack in the discharge leg pipe under static operating pressure is well below the fracture toughness of carbon steel at 550 F operating
- temperature, which has an upper shelf near 250 ksi vin
- Thus, the crack would be stable and would not tend to propagate further under pressure loading alone.
' The crack gpening area for static pressure loacing was calculated to be 1.36 in . Figure 29 shows a magnified view of the crack opening displacements for a one-half circumferential crack under static operating pressure loading. In the dynamic analysis of Section 5.b, the maximum crack opening area under dynamic logding for a one-half circumference crack was determined to be 5.2 in . However, this value was based on the assumption of free (unrestrained) motion at i j both ends cf the pipe. The end restraint provided by the reactor l vessel and discharge pump significantly limits the amount of crack opening which can occur. i E
, ar prw-- -w , , - - . - - - -. ,- ,-----.n,.-- - . . - - , - . - - - - - - - - - - - - - - - - - - - - - -
Attachment to LD-83-053 Pagn 37 6/14/83 c.2 Seismic Loading Conditions In the design of nuclear reactor components for seismic loading, excitations are usually applied in the form of support motion time histories rather than by externally applied forces. The deterministic earthquake response analysis of a structural syst.em must consider these factors which contribute to the input conditions:
- a. simple (rigid-body) translation of the base,
- b. rigid base rotations,
- c. relative motion of different support locations,
- d. the effects of soil-structure interaction where the motion of the base does not directly follow the free-field motion.
For this analysis, all support motions were considered to be the same. This assumption is conservative based on the following method used for determining support time histories. The in-structure response spectrum used to define the SSE loading conditions is shown in Figure 30. The support acceleration time history was generated from the in-structure response spectra using a procedure for generating a seismic artificial time history with a compatible response spectra based on the Fourier transform method (Ref.16,17). An advantage to using an artificial time history is that it can be generated with a short overall duration which maintains the identical design spectrum over the frequency range of interest. The only requirement is that the total duration must be significantly greater than the period of the lowest frequency. For this analysis, a total duration of 6 seconds was chosen for the seismic event, with a rise time of 1 second and a decay time of 2 seconds. The resulting acceleration and velocity time histories for horizontal support motion are shown in Figure 31. In this Figure the maximum c acceleration is 1.2 g from the artificially-generated time history, i and the corresponding peak in velocity is 54 in/sec. Typical values for peak horizontal ground acceleration used in design basis earthquakes are on the order of 0.2 - 0.3 g (Ref. 15). It foilows l from this that the corresponding maximum velocity would be j j approximately 12 in/sec. The seismic loading used in this analysis, ! therefore, represents a "very severe" earthquake. The artificially-generated time history motion was applied to the support locations, and the dynamic response of the discharge leg pipe was evaluated for these seismic loading conditions. The results of l the substructured model, without a crack, were compared with the behavior of the coupled model during seismic loading (Ref.15). The maximum values calculated for the reaction forces and moments at the reactor vessel upper column support are in excellent agreement with the coupled model. This demonstrates the validity of the approach used in the seismic loading of the structure, and verifies the overall dynamic response of the structural model.
Attachment to LD-83-053 Pag 2 38 6/14/83 c.3 Elastic-plastic Dynamic Analysis The natural frequencies of the reactor coolant system were extracted by a modal analysis. It was determined that the first (lowest) natural frequency of the discharge leg pipe without a crack was 16-17 cycles per.second. This is within the range of frequencies which would contribute to normal modes during seismic loading.- The static load state of the system with operating pressure was used as the initial state for tne dynamic analysis. The pipe was considered to be uncracked at time t=0 and a circumferential crack was presumed to initiate at a critical point during the seismic event. A criterion for crack initiation would, for example, be maximum tensile strain at the terminal end of the reactor vessel inlet nozzle. Generally it can be argued that maximum strain is produced in a structure during a seismic event shortly after the peak acceleration.
~
On this basis, a large time step of 0.1 seconds was chosen for the early portion of the dynamic analysis during the " buildup" phase of the earthquake. Direct integration of the dynamic equations was performed using the Newmark-Beta method. Stability problems did not arise since a recycling method was used to ensure that dynamic equilibrium was satisfied at the end of each time step within a set tolerance. In addition, a small amount of ness a ng as ncMed wM a damping factor of 1.0 x 10-5 This damping factor imposes less than
.05 percent damping on modes with frequencies lower than 100 Hz.
From the time history plots in Figure 31, it is apparent that the maximum positive seismic excitation occurs at a time of t=1 second. This corresponds to the most severe externally applied loacing for a one-half circumference crack around the outside of the pipe elbow. Prior to the time of most severe loading, a smaller time step of 0.01 seconds was introduced at time t=0.9 seconds while the pipe remained uncracked. The smaller time step was chosen to delineate the high frequency response of the pipe (16-17 Hz) which would also contribute to the crack opening. Figure 32 shows a response curve of the velocity time history at the midpoint of the discharge leg pipe. " Smoothing" effects due to the l large time steps can be seen for time t < 0.9 seconds, whereas, the
- higher frequency response is apparent for time t > 0.9 seconds when the smaller time steps were used.
( I A determination of the most critical time to release the crack during i the seismic analysis was based on energy principles. On that basis, it was determined that the loadings would produce maximum crack opening for a suddenly appearing crack initiated at t = 0.99 seconds. l t
Attachment to LD-83-053 Pagm 39 6/14/83 A second analysis was performed which included the introduction of a one-half circumference crack at time t = 0.99 seconds. The dynamic behavior of the pipe with the crack was traced with time steps of 0.01 seconds to determine the maximum crack opening. Local plasticity effects were taken into account at the crack tip region for stress levels exceeding the yield strength of the material. Work hardening effects were included for SA-516 Gr-70 carbon steel using the stress-strain curve shown in Figure 21 for this material. Kr values were calculated at each time step using the J-Integral tec5nique for determination of the crack tip stress intensity factor. The analysis was continued for a sufficient number of time steps to determine the total extent of crack opening and the maximum stress intensity at the crack tip due to combined pressure and seismic loading. For comparison with the generated seismic selocities, a plot of the velocity time history at the base of the reactor vessel support column is shown in Figure 33. The analysis was carried out to a total time of 1.19 seconds, well past the peak in the velocity curve. A similar velocity time history at the midpoint of the discharge leg pipe is given in Figure 34 A noticeable difference in the velocity profile occurs at the point of crack opening. The change is apparent in both the magnitude and frequency of the natural periodic motion. A comparison with the uncracked velocity time history shown in Figure 28 indicates that the peak velocity is reduced due to the incidence of the crack. This indicates a reduction in the kinetic energy of the pipe resulting from a change in stiffness. In effect, energy in the pipe is reduced due to crack opening, and the response of the pipe with the crack is substantially different due to the change in stiffness of the system. c.4 Resulting Crack Behavior Crack opening effects are best described in terms of the stress intensity factor, Kg . A plot of Kg vs. time starting at the point of crack initiation is shown in Figure 35. The rapid release of energy following crack initiation is characterized by the sharp increase in Kr to a value of 107 ksi V in.' The increase in Kg due to dynamic effects is approximately 16 percent above the static value of pressure loading alone. The response of the pipe following the opening of the crack causes fluctuations in the value of Kg about the average static value of 92 ksi ViiG The contribution of the seismic loading to the crack opening is small in comparison to the pressure effects. This is due to the fact that the structural stiffness at the reactor vessel and pump enas of the ciscnarge leg pipe severely restricts the rotations and displacements which are a prerequisite to large crack opening effects. The support provided by the reactor vessel and the pump tends to hold the pipe in place so long as a portion of the pipe remains intact. This end constraint severely limits the effects of seismic loading on a crack in the discharge leg pipe. The calculated values for Kg remain well below the critical value for crack instability.
Attachment to LD-83-053 Pagn 40 6/14/83 The maximum crack opening area for a one-half circumference crack under pressure and seismic loadings was calculated to be 1.65 in.2, A 1.65 in.2 crack openin opening area of 5.42 in gwhich can be compared would result from with the maximum crack a one-half circumference crack in a pipe without the end constraint afforced by the reactor vessel and reactor coolant pump. From this work it can be concluded that circumferential cracks must be larger than halfway around the circumference before the effects of both pressure and SSE could cause rapid crack extension.
- d. Axial Slot in Elbow In addition to circumferential or guillotine type breaks, axial cracks or slots are included in the list of design basis pipe breaks.
Reference 1 shows that slots are to be hypothesized on the inside of two of the pump suction leg elbows. In order to evaluate the " leak-before-break" concept and compute leakage areas in the pump suction elbow, a finite element model of one half of a 90* elbow was constructed using shell elements available in the MARC program. The shell model is augmented with a beam at the end to facilitate the application of boundary conditions and the moment loads on the elbow end. Appropriate boundary conditions are chosen on the lines of symmetry and ends. Details of a typical mesh in the surface coordinate system are shown in Figure 36. Figure 37 shows the mesh in a cartesian coordinate system. A number of different crack lengths, 8,12,16 and 20 degrees, on the half elbow are chosen for leakage area and Jr calculations. The crack lengths correspond to approximately 8,12,16 and 20 inches crack length at the inside radius of the torus on the 90* elbow. Operating pressure (2250 psi) and the maximum bending moment due to normal plus seismic loadings (100,000 in-lb) are applied to the cracked elbow. An elastic static analysis was then performed. The total crack opening area and J integral vs crack length are presented in Figure 38. The leakage area ranges whereas J varies from 0 to 500 inib/in.{ rom 0 to 0.7 square inches Using the value of 885 gpm/in.2 of crack area, the 8 inch long crack would be clearly detectable with a leakage of about 60 gpm. A crack of length much less than 8 inches therefore would be detectable at a 10 gpm rate. An elastic plastic analysis of the 20 inch long crack in the pump suction leg elbow was performed to assess the conservatism of the elastic analysis. The result loading condition was 1.0 and in.{ngthe crack correspondingopening Ji was area for the same essentially unchanged. This result indicates that the elastic analysis for leak-before-break evaluation is conservative because the elastic analysis produces a lower leakage area but (for these loadings) essentially the same value of Jy .
Attachment to LD-83-053 Paga 41 6/14/83 The maximum value of the Tearing Modulus for the largest crack . analyzed is: T= 1J E 258 cr; 2
= 50 x 30 x 10 6 = 2.4 3
(26 x 10 )2 An evaluation of the contributions of the pressure and moment loadings shows that only the pressure loading contributes to the stress concentration and leakage areas. The end moment has negligible effect on these parameters because the moment predominantly produces only axial stress in the elbow which has no effect on crack opening. I e 4 1
Attechment to LD-83-053 Pega 42 6/14/83
- 6. MARGINS In order to assess the safety margins for the leak-before-break condition in the main loop piping of System 80 plants, two analyses have been performed. The first analysis is intended to establish the margin of safety on the seismic load carrying capability of the pipe with a leaking crack. The second analysis is intended to establish the margin of safety on the leaking crack length relative to the critical crack length.
- a. Safety Margin on Seismic Load Carrying Capability To assess the margin and seismic load carrying capability, an analysis was performed to determine the likelihood of extension of a crack which is just large enough to leak at a 10 gpm rate during normal operation. The SSE loading was applied and the ductile fracture J-resistance curve was conservatively considered to be degraded to 257. of nominal properties of JIC and Tmat. These properties are shown in Figure 39.
The analysis procedure is essentially the same as that used for the seismic analysis in Section Sc. The finite element mesh in the crack region has been made to be consistent with the mesh used for the static crack opening calculations in Section 4a. The overall system model has been extended to include the steam generators in order to obtain more accurate system response. The finite element mesh is shown in Figure 40. From the crack opening area calculation described in Section 4a, cracks of the same size in the top and outward side of the discharge leg terminal end are found to open essentially the same when subjected to normal operating loads. For a 10 gpm leakage, Figure 17 indicates that a crack length of 7.5 to 8 inches would be required. For the seismic loading analysis, therefore, the crack length is assumed to extend over a 30' arc since that results in a crack length on the inside diameter of the pipe of 7.85 inches. The SSE loading was developed f rom the seismic response spectra by the fourier transform method for generating an artificial time history as described in Section Sc. The response spectra utilized is considerec typical for System 80 plants and is shown in Figure 41 and the l generated support displacement history is shown in Figure 42. A comparison of the spectrum with the enveloping spectrum of the analysis of section Sc (Figure 30) further shows the severity of the envelope used previously. The several displacement cycles in Figure 42 have essentially the same form and magnitude. For this analysis, a displacement which envelopes each cycle is applied as a representative loading condition for all cycles during the seismic event. The representative seismic cycle loading is shown in Figure 42. Consideration is given to the number of cycles which the crack tip experiences in assessing crack stability. The normal full power loadings are statically applied to the finite element model containing the 30* circumferential crack to simulate the initial condition for the dynamic analysis. The J integral value and crack opening area are found to be essentially the same as those given in section da, thereby confirming that modelling changes did not
Attcchment to LD-83-053 Pagm 43 6/14/83 significantly effect the static results. The seismic support motions are then applied and the resulting J and crack opening area histories are shown in Figure 43 and 44. The J integral value is not greatly affected by the seismic loadings, which is consistent with the results of the analysis of section Sc. Similarily, the crack opening area increase due to the seismic loads is very small. The J integ value of J IC (150 in ib/in.from galFigure results 39)~ when
, compared show a considerable to the " deg margin against instability.
- b. Safety Margin on Crack Length To assess the margin on crack length, the smallest length of crack which will be unstable when subjected to the SSE loading and evaluated using only 25% of the ductile crack resistance properties must be determined. For this analysis the loading conditions are applied in the same manner as in the analysis of Section 6a. The assumed crack size, however, is much larger. Previous analyses indicated that a crack must be more than halfway around the circumference in order to be unstable even during seismic loading. Those analyses, however, did not consider the significant property reduction employed in this analysis.
The ductile crack stability criterion states that a crac.k will be unstable if: J > J IC and the tearing modulus T = )J d >T mat. g avo-' In order to determine T, more than one crack size must be evaluated so that the derivative term dJ/@ a can be evaluated. It is appropriate, therefore, to consider the cracks for evaluation to be a half circumference crack and a crack somewhat less than half circumference long. Evaluation of these cracks will enable a comparison with previous work and enable an interpolation or slight ! extrapolation to the length which will be unstable for the imposed loads and properties. l The crack opening area results for the normal loadings of section 4a show that the opening is greater for cracks on the top and outward side of the elbow and less for the bottom and inward side. It is assumed, therefore, that the top and outward side are the crack i locations to consider for the most conservative stability evaluation. The two long cracks which are evaluated are a half circumferential crack centered 30* toward the outside from the top and a one third circumferential crack also centered at 30 toward the outside from the i top of the elbow. These assumed cracks are shown in Figure 45. The normal operation loads are statically imposed on the finite element model containing the larger cracks. For th circumferential crack, the j value is 374 ,inatlb/in.g half the "a" end crack tip and 227 in lb/in. at the "b" end crack tip. The average l
Attachment to LD-83-053 Page 44 6/14/83 value is slightly higher than the value found in Section Sc. (The Kg value average of 282ofin 92 ksi y).found Ib/in. This isinreasonable section Sc because corresponds the skewed to a J crack is expected to be slightly more severe than an outward side crack. The excellent agreement, however, confirms the new modelling The J value and gives confidence in the consistency of the results2 at the "a" j for the one third circumferenc end crack tip and 84 in lb/in.gatcrack the "b" is end. 109 in lb/in. A plot of static normal operation load J vs half crack length is shown in Figure 46. For the larger crack sizes the J values fit a cubic curve in crack length very precisely. This cubic relationship enables a more precise calculation of Tapplied. This cubic form was used in section 5b in order to obtain T Pli d ff05 "IY 0"' analysis. AlsoshownisthedegradedJfCwhchindicatesthatfor J The jj IC. cracks value ofgreater than 150* in-circumference, "a" tip ofEhe,Na>lf Jhowever,circumference for the T crack is o$hj b which is much lower than the degraded value of Figure 39. An instability diagram using the degraded values of Figure i 39 is shown in Figure 47. The applied loadings for the normal operation load for the 180* and 120* circumferential cracks in also shown. The margin between the loading values and the instability line show that instability will not occur for these cracks during normal operation even if the degraded properties are considered. In order to compute J for normal operation plus seismic loads the i seismic support displacements are applied in the same manner as in the . previous section and the dynamic analysis procedes in the same way. The resulting J vs time curves for each. of the crack tips is shown in Figure 48 for the 120* crack and in Figure 49 for the 180* crack. The increase in crack .,pening area caused by the seismic loading for both crack sizes is very small compared to ' normal operating conditions of 0.5 in.ghe and 1.6areas in. for detgrminedthe 120* during circumferential and 180* circumferential cracks respectively. The J 11 d and T jj
! not c$00ge8 signif!ESnth from values the normal during operation the seismic loading loading results.are This is consistent with the results of section Sc where only a small
? increase in crack tip loading occurred for a much higher seismic loading. The J ap instabilitydiahpjj,fandT am o Figure bied values,the represent shown seismic on as thewell as the normal operating loading case. This analysis shows clearly, that very long cracks in the main loop piping will remain stable when subjected to normal operation and SSE loads even if the ductile crack resistance material properties are considered to be only a fourth of the nominal value. These results satisfy, with considerable margin, the requirements for the demonstration of leak-before-break in the main loop piping. 4 e - v- ,m -w..,, - , e, ,.,---e .- - - - . - - - , - - - , . . - - - . . . - - - - - - - - , , , - - , - - . . - - - - . + - , -
Atttchment to LD-83-053 Paga 45 6/14/83
- c. Margin on Axial Slot in Elbow
- The results of the analysis of the 20 inch long axial slot in the 90' pipe elbow described in Section 5d, are plotted on the instability diagram in Figure 50. The stability of this crack is evident from the curve, which reflects the very low value of Tapplied*
- 7. CONCLUSIONS In this report', a variety of crack sizes, loadings and material properties have been considered and analyzed in order to demonstrate that a leak-before-break condition exists in the main loop piping of a CE PWR. It was shown that cracks in this piping system would tend to grow through the pipe wall rather than circumferentially due to the fatigue loading conditions.
Conservative crack opening area calculations showed that a circumferential crack about 8 inches long in the top of side of the discharge leg terminal end would leak at a rate of 10 gpm but would have a high margin against instability even when subject to safe shutdown seismic loads. At the bottom of the terminal ends, a greater crack length is required for a leak rate of 10 gpm, but the loading in this region is so small that crack instability is not a concern. The margin against instability for detectably leaking cracks is demonstrated to be substantial in terms of margin on material propertes, on loadings, and on crack size required for instability. An axial slot less than 8 inches.in length was shown to leak more than 10 gpm. This size crack has a very high margin against instability since even 20 inch long cracks are clearly stable. The requirements for the demonstration of leak-before-break therefore are also satisfied for axial slot cracks in the piping elbows. 1 l
, , + - e s. n , - - , - -
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Attachment to LD-83-053 Page 46 6/14/83
- 8. REFERENCES
- 1. " Design Basis Pipe Breaks for the Combustion Engineering Two Loop Reactor Coolant System," CENPD-168-A, Combustion Engineering, June 1977.
- 2. Reference Deleted
- 3. " Crack Size Considerations in Primary Piping," CENPD-78, Combustion Engineering, December 1973, also Appendix C of Reference 1.
- 4. Griesbach, T. J., " Dynamic Elastic Plastic Behavior of Circumferential Cracks in a Pipe Subjected to Seismic Loading Conditions", Presented l at 1980 Pressure Vessels and Piping Technology Conference, San l Francisco, 1980.
- 5. ASME Boiler and Pressure Vessel Code Section XI, Article A-4000,
" Definition of Material Properties", New York,1972.
- 6. %res, D. J., Griesbach, T. J., DeSaulniers, W. E., "An Evaluation of Leakage for Postulated Circumferential Cracks in PiR Primary System Piping", Paper F 7/4 6th SMiRT, Paris, France,1981
- 7. Peck, D. A. " Determination of the Appropriate Stable Crack Size to be Used in A Rational Design Basis for Pipe Breaks", Paper 7/2, 6th SMiRT, Pari s, France,1981.
- 8. USNRC Regulatory Guide 1.45 " Reactor Coolant Pressure Boundary Leakage Detection Systems", May,1973.
l
- 9. Mayfield, M. E., et al " Cold Leg Integrity Evaluation", NUREG/CR 1319, prepared by Battelle Columbus Laboratories for the US NRC, Washington D.C., February,1980.
- 10. Henry, R. E., and Fauske, H. K., "The Two-Phase Critical Flow of One-Component Mixtures in Nozzles, Orifices and Short Tubes", Journal of Heat Transfer, Vol. 93, pp.179-187,1971.
l 11. Gudas, J. P., " Piping Material J R Curve Characterization", presented at HSST Review Meetingf July 24, 1980, David Taylor Naval Ship R+0 Center, Annapolis, M).
- 12. Parks, D. M., "A Stiffness Derivative Finite Element Technique for Determination of Elastic Crack Tip Stress Integrity Factors,"
International Journal of Fracture, Vol.10, No. 4, December,1974.
- 13. Ayres, D. J., " Determination of the largest Stable Suddenly Appearing Axial and Circumferential Through Cracks in Ductile Pressurized Pipe",
Paper F 7/1, 4th International Conference on Structural Mechanics in Reactor Technology, San Francisco, August,1977.
.,n>,--, -
Attachment to LD-83-053 Pags 47 6/14/83
- 14. Griesbach, T. J., and Ayres, D. J., " Opening and Extension of Circumferential Cracks in a Pipe Subject to Dynamic loads", Paper F 5/1, 5th International Conference on Structural Mechanics in Reactor Technology, Berlin (West), Germany, August,1979; to be published in Nuclear Engineering and Design .
- 15. Gerdes, L. D., " Dynamic Structural Analysis of Uncoupled Subsystems",
Paper K 6/18' 4th International Conference on Structural Mechanics in Reactor Technology, San Francisco, August, 1977.
- 16. Scanlon; R. H.,' and Sachs, K., " Earthquake Time ' Histories and Response Spectra"; Journal of Engineering Mechanics Division, ASCE, Volume
~
100, August, 1974, pp. 635-655.
- 17. Scanlan, R. H., and Sachs, K., " Floor Response Spectra for Multi-l Degree-of-Freedom Systems by Fourier Transform", Paper K 5/5, 3rd International Conference on Structural Mechanics in Reactor Technology, London, U.K., September, 1975.
F45393 l t i
Attachment to LD-83-053 Page 48 6/14/83 10-2 , 10-3 _ DATA 5:
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Attachment to LD-83-053 Page 53 6/14/83 HOT LEG TERMINAL END .
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Attachment to LD-83-053 Page 99 6/14/83 Table 1 LOADING TRANSIENTS ANALYZED AND LIFE OCCURENCES Life Loading transients Occurances
- a. Plant heatup,1000F/hr 500
- b. Plant cooldown,1000 F/hr 500
- c. Plant loading, 5%/ min. 15,000
- d. Plant unloading, 5%/ min. 15,000
- e. 10% step load increase 2,000
- f. 10% step load decrease 2,000
~ 6 10
- g. Normalplantvarjation
(! 100 psi, t 10 F) Reactor Trip 400 h. 0 200
- 1. Leak test, 2250 psia,100 F 400 F U 10
- j. Hydrostatic test, 3125 psia,100 F-400 F 40
- k. Loss of Reactor Coolant Flow (*) ,
40
- 1. Loss of Turbine Generator Load (*)
Loss of Secondary Pressure (*) 5 m.
- n. Operating Basis Earthquake (*) 200
- Abnamal Transient Conditions
Table 2 LOADS AND ELASTIC MATERIAL PROPERTIES USED FOR CRACK OPENING AREA STRUCTURAL ANALYSES COMBINED NORMAL OPERATING LOADS FOR THE'RMAL EXPANSION, WEIGHT, AND 2260 PSIA INTERNAL PRESSURE FX Fy Fz MX MY My LBS LBS LBS FT-LBS FT LBS FT LBS I 1 6 6 5 0 5 4 -1.719x10 2.518x10 , EL80W -8.570x10 1.204x10 1.157x10 -2.619x10 DLTE 6 6 5 6 5 NOZZLE -1.202x10 -9.167x10 -1.533x10 9.103x10 -1.976x10 -3.414x10
-6.406,t10 6 5 6 m HLTE PIPE -7.509x10 0.0 3',071x10 0,'O 0.0
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- v0uNG'S MODULUS = 2.900 m 107 PSI g POISSON'S AATSO = 0.290 C N
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