ML17309A257
| ML17309A257 | |
| Person / Time | |
|---|---|
| Site: | Ginna  |
| Issue date: | 04/26/1982 |
| From: | Maier J ROCHESTER GAS & ELECTRIC CORP. |
| To: | Crutchfield D Office of Nuclear Reactor Regulation |
| Shared Package | |
| ML17256A855 | List: |
| References | |
| NUDOCS 8204280230 | |
| Download: ML17309A257 (428) | |
Text
Steam Generator Evaluation Ginna Steam -Generator=-Tube=Failure.Tncident.
January 25,. 1982 R. E. Ginna Nuclear Power Plant Docket No. 50-244 gNll@M7. IIIIIlIBI:II,f. IlIIPE April 26, l982
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TABLE OF CONTENTS PAGE
1.0 INTRODUCTION
1.1 General 1.1-1 1.2 Inspections 1~2 1 1.3 Repairs 1 3 1
~
1.3.1 Phase I 1.3-1 1.3.2 Phase II 1.3.-1 1.4 Failure Analysis 1.4-1 1.5 Post Repair Program 1.5-1 1.6 Conclusions 1.6-1 1.6.1 Initial Plugging 1.6-1 1.6.2 Subsequent Degradation 1.6-1 1.6.3 Tube Rupture 1.6-2 1.6.4 Adequacy of Repairs 1.6-2 2.0 STEAM GENERATOR CONFIGURATION AND OPERATING HISTORY 2.1 Configuration 2.1-1 2.2 Operating History 2.2-1 2.2.1 Chemistry 2 2 1
~
2.2.2 Plugging 2.2-2 2.2.3 Secondary Side Modifications 2 ~2 3 3.0 STEAM GENERATOR INSPECTION RESULTS 3.1 Eddy Current Examination 3a1 1 3.2 Profilometry Examination 3.2-1 3.3 Fiber Optics Inspection 3 ~3 1 3.4 Television Video Inspection 3.4-1 3.5 Foreign Objects 3.5-1 3.5.1 A-Steam Generator 3.5-1 3.5.2 B-Steam Generator 3.5-1 4.0 STEAM GENERATOR TUBE FAILURE ANALYSIS PROGRAM 4.1 Introduction 4.1-1 4.1.1 General 4.1-1 4.1.2 Program 4.1>>1 RECOMTO'5 OOCKET FILE COPY
TABLE OF CONTENTS (continued)
PAGE 4.2 Postulated Failure Mechanisms 4.2-1 4.2.1 Initial Plugging 4.2-1 4.2.2 Collapse 4.2-1 4.2.3 Severance 4.2-1 4.2.4 Wear 4.2-2 4.2.5 Summary 4.2-2 4.3 Data Review 4.3-1 4.3.1 Purpose 4. 3-1 4.3.2 Examination and Repair History 4.3-1 4.3.3 Evaluation 4.3-4 Metallurgical Review 4.4-1 4.5 Analysis 4.5-1 4.5.1 Design Parameters 4.5>>1 4.5.2 Thermal Hydraulic Evaluation 4.5-4 4.5.3 Axial Loads 4.5-9 4.5.4 Lateral Impact Loads 4.5-12 4.5.5 Collapse 4.5<<15 4.5.6 Flow Induced Vibration 4.5-18 4.5.7 Fatigue 4.5-22 4.5.8 Wear 4.5-26 4.6 Model Tests 4.6-1 4.6.1 Test Objectives 4.6<<1 4.6.2 Test Apparatus 4.6-1 4.6.3 Test Procedure and Results 4.6-2 4.6.4 Conclusions 4.6-4 4.7 Laboratory Tests 4. 7-1 4.7.1 Collapse Test 4.7-1 4.7.2 Fatigue Test 4.7-2 4.8 Conclusions 4.8-1 5.0 STEAM GENERATOR REPAIR PROGRAM 5.1 Access Holes 5.1-1 5.1.1 Description 5.1-1 5.1.2 Installation 5.1-1 5.1.3 Design 5.1<<1 5.2 Tube Removal 5.2-1
=
5.2.1 Summary 5.2-1 5.2.2 Categorization of Tubes 5.2-1 5.2.3 Tube Removal 5.2-2 5.2.4 Severed Tube Removal 5.2-2 5.2.5 Tube Pull 5.2-2 5.2.6 U-Bend Restraint 5.2-2 5.2.7 Remaining Tubes 5.2-2
TABLE OF CONTENTS (continued)
PAGE 5.3 Loose Parts Removal 5.3-1 5.4 Mechanical Plug Removal 5.4-1 5.5 Material Control 5.5-1 5.6 Post Repair Inspections/Tests 5.6-1 5.7 Radiation Exposure 5.7-1 5.7.1 Planning '5.7-1 5.7.2 Tube Removal 5.7-1 5.7.3 Exposure 5.7-1 6.0 TECHNICAL BASIS FOR REPAIRS 6.1 Introduction 6.1-1 6.2 Analyses 6.2-1 6.2.1 Access Ports 6.2-1 6.2.2 Thermal/Hydraulic Evaluation 6.2-2 6.2.3 Structural Evaluation 6.2-4
' 7.0 FUTURE ACTIVITIES 7.1 Metallurgical Examination 7~1 1 7.2 Testing 7~2 1 7.2.1 Westinghouse 7~2 1 7.2.2 Combustion Engineering 7~2 1 7.3 Loose Parts Monitoring System 7~3 1 7.3.1 Summary 7 ~3 1 7.3.2 Description 7 ~3 1 7.4 Intermediate Outage 7.4-1 Appendix A Metallurgical Examination of Ginna Steam Generator Tubes Appendix B EPRI/CE Analyses and Tests, Steam Generator Tubing Structural Analysis Appendix C Ginna Steam Generator External Tube Loading Test
LIST OF TABLES TABLE NO. TITLE 2.1 A-Steam Generator Tube Plugging History 2.2 B-Steam Generator Tube Plugging History 2.3 B-Steam Generator Periphery Area Defects 3.4-1 Data Review B-Steam Generator Hot Leg 4.3-1 Number 6 Wedge Area Plugged Tubes - Inspection History 4.3-2 Number 4 Wedge Area Plugged Tubes - Inspection History 4.5.2-1 Nominal and Between Tube Crossflow Velocities 4.5.3-1 Summary of Worst Case Axial Loads on a Plugged Tube 4.5.4-1 Foreign Object Induced Loads 4.5.6-1 Basic Analysis Model Geometry 4.5.6>>2 Tube Fundamental Frequencies for the Various Cases Analyzed 4.5.6-3 Vortex-Shedding Induced Loads on Protrusions for a Fluid Velocity of W.O. Ft./Sec.
4.5.7-1 Calculated Stresses (KSI) - Nominally Plugged Tube 4.5.7-2 Principal Stresses (KSI) - Nominally Plugged Tube 4.5.7-3 Usage Factor for a Nominally Plugged Tube 4.5.7-4 Usage Factor for a Nominally Plugged Tube With Notch or Stress Riser 4.5.7-5 High Cycle Fatigue Curve 4.5.7-6 High Cycle Fatigue Curve 4.8-1 Conclusions about Collapse 4.8-2 Conclusions about Fatigue 4.8-3 Conclusions about Wear 4.8-4 Conclusions about Burst
'l f
0
LIST OF TABLES TABLE NO. TITLE 5.1 B-Steam Generator Recovery Exposure 6.2.1-1 Stress Summary for 3 Inch diameter Access Port, 6.2.2-1 Between Tube Crossflow Velocities in and Near the Tubes Removed Region 6.2.3-1 Summary of Maximum Tube Gap Velocities 6.2.3-2 Summary of Vortex Shedding & Turbulence Analyses
LIST OF FIGURES FIGURE NO. TITLE 2.1 Series 44 Steam Generator 2.2 Wedge Area Configuration 2.3, A-Steam Generator Plugged Tube Map 2.4 B-Steam Generator Plugged Tube Map
'3.4-1 Secondary Side Tube Sheet Periphery Video
- 3. 4-2 B-Steam Generator Hot Leg 4.2-1 Postulated Failure Mechanism Sequence 4.3-1 Number 4 Wedge Area: Plugging History 4.5.1-1 Schematic Showing Hot Leg Span Geometry 4.5.2-1 Charm Model 4.5.2-2 Three Dimensional (WECAN) Hydraulic Model of Tube sheet to First Support Plate Region 4.5.2-3 R-2 Distribution of Nodes for 3-D Hydraulic Analysis 4.5.2-4 R-9 Distribution of Nodes for .3-D Hydraulic Analysis 4.5.2-5 Computed Average Velocities 4.5.2-6 Base Case Computed Average Velocities 4.5.2-7 Velocity Vector and Quality Distributions 4.5.2-8 Velocity Vector and Quality Distributions 4.5.2-9 Velocity Vector and Quality Distributions 4.5.2-10 Velocity Vector and Quality Distributions 4.5.2-11 Fluid Velocities in the Node 11 Cell 4.5.4-1 Dynamic Model of Foreign Object Impact 4.5.4-2 Steam Generator Tube Spring Force Time History - Case 1
LIST OF FIGURES FIGURE NO. TITLE 4.5.4-3 Foreign Object and Steam Generator Tube Displacement Time History - Case 1 4.5.4-4 Steam Generator Tube Spring Force Time History - Case 2 4.5.4-5 Foreign Object and Steam Generator Tube Displacement Time History Case 2 4.5.4-6 Steam Generator Tube Spring Force Time History - Case 3 4.5.4-7 Foreign Object and Steam Generator Tube Displacement Time History - Case 3 4.5.5-1 External Pressure and Axial Load Required for Incident Yielding
- 4. 5. 5-2 Computer Model for Uniform Pressure Loading 4.5.5-3 Model Used for Concentrated Loading 4.5.5-4 Boundary Conditions for Concentrated Load Model 4.5.5-5 Concentrated Radial Loads 4.5.5-6 Axial Distance from Load 4.5.6-1 Basic Analysis Model Geometry 4.5.6-2 Fluidelastic Stability Ratio of Tubes-Fixed-Fixed 4.5.6-3 Fluidelastic Stability Ratio of Tubes-Fixed-Pinned 4.5.6-4 Fluidelastic Stability Ratio Comparison 4.5.6-5 Cylindrical Cross Section Tube - Cross Flow Velocity 4.5.6-6 Flat Cross Section Tube Cross Flow Velocity 4.5.6-7 Tear Model for Calculating Illustration of Alternating Lift and Frequency vs. Fluid Velocity for a Tear
LIST OF FIGURES FIGURE NO. TITLE 4.5.6-9 Illustration of Alternating Lift and Frequency vs. Fluid Velocity for a Tear 4.5.6-10 Illustration of Alternating Lift and Frequency vs. Fluid Velocity for a Tear 4.6.2-1 Plan view of Flow Test Model 4.6.2-2 Section View of Flow Test Model 4.6.2-3 Rear Elevation View of Flow Test Model 4.6.2-4 Photograph of Tube Bundle Used in Flow Test Moel 4.6.2-5 Region of Test Simulation 4.6.2-6 Test Simulation Detail 4.6.2-7 Cold Flow Loop 4.6.2-8 .Biaxial Accelerometer Orientation 4.6.2-9 Force Transducer Orientation 8 Installation Assembly 4.6.2-10 Photograph of Force Transducer Mounted in Tube Support Plate
- 4. 6.3-1 Foreign Object
- 4. 6.3-2 Tube Accelerometer Response Time Histories for Foreign Object Impact 4.6.3-3 Tube Force Transducer Time Histories for Foreign Object Impact
- 4. 6.3-4 Tube Response for Foreign Object Impact
- 4. 6.3-5 Tube Accelerometer Response Time Histories for Instrumented Impact Hammer 4.6.3-6 Tube Force Transducer Response for Instrumented Hammer Impact 4.6.3-7 Accelerometer Calibration Data
.4.6.3<<8 Force Transducer Calibration Data
LIST OF FIGURES FIGURE NO. TITLE 4.6.3-9 Tube Wear Photograph 4.6.3-10 Tube Wear Photographs 4.6.3-11 Tube Accelerometer Response Envelopes 4.6.3-12 Tube Accelerometer Response with Tube Degradation 4.6.3-13 Severed Tube Accelerometer Response 4.7.1-1 Test Fixture for, Collapse of Steam Generator Tubes 4.7.1-2 Photograph Showing the Test Set Up for Tube Collapse Test 4.7.1-3 Photograph Showing Instrumentation Details for the Tube Collapse Test 4.7.2-1 View of Test Set Up 4.7.2-2 View of Test Set Up 4.7.2-3 Close View of Fatigue Test Set Up 4.7;2-4 Close View of Fatigue Test Set Up.
4.7.2-5 Instrumentation Block Diagram 4.8-1 Postulated Failure Mechanism Sequence 5.1 inch access hole cross sectional view I'hree 5.2 Number 4 Wedge Area 5.3 Number 6 Wedge Area P
5.4 Access Hole Cover Plate Assembly Details 5.5 Categorization o f De fects, B-Steam Generator 6.2.1-1 Location of Access Ports 6.2.1-2 Finite Element Model 6.2.2-1 Charm Lateral and Axial Velocity Values 6.2.2-2 Charm Lateral and Axial Velocity Values
LIST OF FIGURES FIGURE NO. TITLE 6.2. 2-3 Charm Lateral and Axial Velocity Values 6.2.2-4 Charm Lateral and Axial Velocity Values 6.2.2<<5 Quality Distribution 6.2.2-6 Quality Distribution 6.2.3-1 Collapse Pressure of Tubing with Elliptical Wastage 6.2.3-2 Collapse Pressure of Axially Stolled Tubing 6.2.3-3 Collapse Pressure of Uniformly Thinned Tubing 6.2.3-4 Schematic of a Partial Tube 7.1 Steam Generator Loose Part Monitor Sensor Locations 7.2 Electrical One Line Steam Generator Loose Part Detection System B.1 Worst Load Conditions Prior to Tube Plugging B.2 Worst Load Conditions After Tube Plugging B.3 First Mode Natural Frequency as a Function of Axial Compressive Load B.4 Fluid-Elastic Vibrations B.5 Temperature Distributions for Various Loading Conditions B.6 Computer Print Pattern of the Element Layout B.7 Element Layout B.8 Temperature vs. Time Curves B.9 Downcomer Temperature Distribution for Hot Standby B.10 Downcomer Temperature Distribution for 100%
Power Steady State B.ll Downcomer Temperature Distribution for Cold Feed at Hot Standby
~ B.12 Transient Vs. Time Downcomer Annulus Plot Outlet Temperature
LIST OF FIGURES FIGURE NO. TITLE B.13 Mass Quality at 7 Inches Above Tube Sheet B.14 Radial Velocity at 7 Inches Above Tube Sheet B.15 Axial Velocity at 14 Inches Above Tube Sheet B.16 Void Fraction at 7 Inches Above Tube Sheet B.17 Mass Quality at 20 Inches Above Tube Sheet B.18 Radial Velocity at 20 Inches Above Tube Sheet B.19 Axial Velocity at 27 Inches Above Tube Sheet B.20 Mass Quality at.33 Inches Above Tube Sheet B.21 Radial Velocity at 33 Inches Above Tube Sheet B.22 Mass Quality at 91 Inches Above Tube Sheet B.23 Radial Velocity at 46 Inches Above Tube Sheet 0 s.~~ Axial Velocity at 51.8 Inches Above Tube Sheet B.25 Mass Quality at 46 Inches Above Tube Sheet B.26 Radial Velocity at 91 Inches Above Tube Sheet C.1 External Tube Loading Test Apparatus
Steam Generator Evaluation
1.0 INTRODUCTION
1.1 General At 9:25 a.m. on January 25, 1982, a tube ruptured in the B-Steam Generator at Ginna Station. A report summarizing the sequence of events, operator actions, emergency procedures, equipment performance, radiological assessment, and recommendations relative to the transient itself was submitted to the NRC by letter dated April 13, 1982. This report provides information specifically related to the tube failure. The work described in this report is a culmination of work performed by several organizations including Rochester Gas and Electric, Westinghouse, EPRI, and Combustion Engineering.
Shortly following the achievement of cold shutdown conditions on January 26, 1982, a comprehensive corrective action program was developed for the B-Steam Generator. The objectives established for this program were:
a) determine the full extent of defects, b) determine the tube failure mechanism(s),
c) restore the steam generator to a condition which is safe to operate while maintaining radiation exposures as low as reasonably achievable, and d) -obtain NRC concurrence for return to power.
To accomplish these objectives, an extensive series of inspections, tests, repairs, and analyses have been performed.
1.1-1
The elevation of the tube rupture was determined by filling the secon'dary side of the steam generator with water, and observing the wide range level as the generator was drained. The level dropped to below the bottom level tap before stopping. Since the tap is centered 12 inches above the tube sheet, this meant, that the rupture was at the lower end of the tube. Observation of the tube sheet from the primary side manways as a small amount of water was added showed water leaking from the inlet, or hot leg, side of tube Row 42 Column 55 (R42 C55). Subsequent fiber optic inspection of the inside diameter (I.D.) of the tube confirmed the presence of the burst.
Extensive eddy current, fiber optics, video, and visual inspections have been performed in both the A and B steam generators. Eddy current examinations have been performed in 100% of the hot, leg tubes and a random sample of cold leg tubes in each steam generator.
Profilometry examination was performed in a random sample of tubes in the B-Steam Generator. A combination of fiber optic and video inspections were performed on the secondary side of each steam generator at the tube sheet. These inspections included viewing the tube bundles from the periphery of both the hot and cold legs, and from the tube lane in the B-Steam Generator.
Additional fiber optic and video inspections of the Number 4 and Number 6 wedge areas in the B-Steam Generator were performed through access holes drilled in the shell of the steam generator.
Visual inspections of the upper internals or steam drum of each steam generator were also performed.
1.2-1
1.3 R~e airs 1.3.1 Phase I The repairs to the B-Steam Generator were performed in two phases.
Phase I included plugging of R42 C55; drilling two, 3 inch diameter access holes in the secondary shell; and removing metallurgical samples of selected tubes. The removal of= metallurgical samples involved the cutting of several sections of tubing from the Number 4 wedge area; and pulling the hot leg side of R45 C47 from the top of the tube bundle.. The results of the metallurgical examination of these tubes are reported in Appendix A to this report.
1.3.2 Phase II The Phase II repairs included removal of structurally degraded tube sections from between the tube sheet and first support plate, and removal of mechanical plugs from 3 acceptable tubes surrounding R42 C55. These 3 tubes had been preventatively plugged prior to performing the secondary side video inspections. Metallurgical examination of the tube sections removed during this phase is presently in progress. Tube ends were restrained in the U-bend region where sections of tubing had been removed. Foreign objects were removed from both steam generators; and tubing fragments were removed from the B-Steam Generator during the repairs.
1~3
1.4 Failure Anal sis The failure analysis program has consisted of metallurgical examinations, analyses, and testing. The metallurgical examinations included visual inspections, photography, radiography, metallography, and scanning electron microscopy. Analyses have been performed to quantify the potential effect of lateral loads, flow induced vibration, local fluid loads, and axial loads on tube failure.
Laboratory testing is being performed to demonstrate the effect of lateral loads, fatigue, and axial loads on steam generator tubing. Model testing has been performed to study the behavior and interaction of foreign objects and tubing in a simulated flow environment.
l 1.5 Post Re air Pro ram A series of inspections and tests will be performed in the B-Steam Generator following completion of repairs. This will include eddy current examination of tubing adjacent to any area involved in the repairs, another secondary side video inspection at the tube sheet, and primary and secondary side hydrostatic tests. A loose parts monitoring system will be installed on the primary and secondary side of each steam generator prior to start-up. An intermediate outage is scheduled to perform another set of steam generator inspections.
1.5-1
1.6 Conclusions 1.6.1 Initial Plu in The eddy current examinations of periphery tubes in the B-Steam Generator showed both I.D. and O.D. defects. The metallurgical examinations of the tube pulled in 1978 showed evidence of O.D.
surface cold working and an O.D. bulge. The metallurgical examinations of the periphery tube sections removed from the steam generator during the present outage have also found cold working of O.D. surfaces. These examinations have also found O.D. surface defects. None of the examinations to date has found any evidence of corrosion or I.D. defects.
Axial load analysis shows that there is insufficient force developed to initiate the tube damage. Model testing has demonstrated that foreign objects are very mobile in the flow field at the bottom of the generator. This testing also showed that the flow velocities are high enough to cause relatively large foreign object impact loads on steam generator tubes. The magnitude of the loads has been corroborated by analysis. Therefore, the most probable cause of the initial tube plugging appears to have been defects resulting from foreign object impacts. However, axial load may have been a contributing factor in the initial plugging of some tubes.
1.6.2 Subse ent De radation Video and subsequent metallurgical examination has shown varying degrees of degradation of previously plugged tubes. In addition to minor dings and similar small defects, this degradation included collapse, severing, and wear. The structural degradation of tubes (collapse," severing, or through wall wear) occurred only in the Number 4 and Number 6 wedge areas. Analysis and test results show that external operating pressures alone will not collapse a plugged tube. The results show that weakening of the tube by wall thinning or ovalization, and/or additional stress from lateral or axial loads, must exist in combination with external pressure, to cause collapse. The metallurgical examinations found significant cold working of the collapse surfaces.
Model testing and calculations have been performed which show that a foreign object can impose the magnitude of lateral loads required to contribute to collapse or severing. The fatigue analysis and testing show that a collapsed tube under sufficient lateral and axial load will fail (sever). Axial load analysis shows that, assuming lockup at the first support plate, axial loads are significantly higher in wedge area tubes. Severing of collapsed tubes in the Number 4 wedge area, but not the Number 6 wedge area, appears to be related to a foreign object residing there for a longer period of time. Vibration analysis shows that collapsed tubes are subject to flow induced vibration, and that severed tubes vibrate with an amplitude sufficient to interact with adjacent tubes.
- 1. 6>>1
Review of the eddy current data'nd metallurgical examination results for tube R42 C55 shows that it went from less than 40% to 84% wall thinning between Nay 1981 and January 25, 1982. Wear calculations for one Inconel tube rubbing on another are consistent with this rate. The metallurgical examinations show that R42 C55 had a ductile, tensile overload failure. Burst strength calculations, assuming nominal ultimate strength for Inconel tubing, show that.
0.875 inch O.D. tubes will burst at 87/ uniform, circumferential wall thinning. The metallurgical examinations show no evidence of corrosion or surface irregularities on R42 C55. All of the available evidence shows that tube R42 C55 ruptured due to excessive hoop stress from internal pressure. The excessive hoop stress was caused by thinning of the tube wall from wear by an adjacent tube. The adjacent tube was able to vibrate against R42 C55 since above.
it had severed as a result of the failure sequence described 1.6.4 Ade ac of Re airs All structurally degraded tube sections have been removed from the B-Steam Generator. An extensive program of inspection and removal has been performed to assure that no loose parts remain in the generator. The inspection, analysis, and test results show that the structural degradation of tubes. which led to the rupture of R42 C55 was caused by the presence of a relatively large foreign object in a hot leg wedge area. The results indicate that a foreign object will not cause structural degradation of tubes outside the wedge areas of the steam generator. Removal of the foreign objects will preclude additional structural degradation of active or plugged tubes, and potential rupture of an active tube. In addition, the analyses and tests show that, in the absence of a foreign object, the plugged tubes with minor O.D.
defects will not experience further degradation.
1.6-2
2.0 STEAM GENERATOR CONFIGURATION AND OPERATING HISTORY IQ 2.1 ~E' Ginna Station's steam generators are Westinghouse Series 44 vertical shell and U-tube units of the recirculating design and operation. They are rated at 3,130,000 lbs/hr. steam flow at 725 psig. The steam generator tubing is mil-annealed Inconel 600 conforming to ASTM Specification SB-163-61T. The 3260 tubes are partially rolled into the tube sheet and seal-welded, leaving an approximately 19 inch crevice in the 22 inch tube sheet. The support plates are carbon steel of the drill hole design. Attachment of the support plates to the wrapper and shell is accomplished by wedging the support plate against the wrapper and welding the wedges in place at six equally spaced locations around the periphery of the support plate. There are twelve (12) locations around the circumference of the support plate where wedges could be applied.
However, by design, only six (6) of these locations are used for attachment of each support plate. The support plate in these locations is free of the 0.375 inch diameter flow holes for several rows into the bundle. These areas of the support plate are called "wedge areas". See Fig. 2.1 and 2.2 for details of the configuration.
- 2. 1-1
IV 2.2 0 eratin Histo 2.2.1 ~h' When Ginna Station began hot functional testing in November 1969, the secondary water chemistry control recommended by Westinghouse Electric Corporation was phosphate buffering control. From start-up through 1970, emphasis was placed on pH control only.
As a result, the Marcy/Halstead ratio of sodium to phosphate (NaP04) generally ranged between 2.8 to 6.0. In 1971 through 1972, recommended chemistry was modified to maintain the Marcy/
Halstead ratio to less than 2.6. Early in 1973, when stress corrosion cracking (SCC) had become evident at several plants, a further reduction of the Marcy/Halstead ratio to 2.1 was recommended.
However, as Na/PO4 was concentrating in the sludge at the lower ratios, an acidic condition was being formed which resulted in wastage attack of the tubing. Therefore, a new Marcy/Halstead ratio range of 2.3 to 2.6 was recommended in late 1973.
During the period of phosphate type control, steam generator blowdown rate was a maximum of 16 gallons per minute per steam generator. Other coordinated phosphate control parameters were as follows:
pH 8.5 to 10.6 PO 10 to 80 ppm Free OH 0 Chloride under 75 ppm Since the phosphate buffering chemistry control was difficult to maintain without concentrating an aggressive environment in the sludge-saturation zone of the tube sheet, Westinghouse recommended in September 1974 that all their operating plants switch from phosphate control to the all volatile treatment (AVT) chemistry control. Ginna had experienced some acidic wastage type attack which was identified by an eddy, current examination in the spring of 1974. In November 1974, Ginna was shut down for steam generator inspection, water lancing, and conversion of water chemistry control to AVT. Steam generator blowdown was increased to 64 gallons per minute per steam generator to assure that cation conductivity would be as low as achievable.
During subsequent operation, cation conductivity progressively improved, decreasing from 2.5 pmhos in December 1974 to approximately 0.7 pmhos by 1977. Typical 1977 blowdown chemistry was as follows:
cation conductivity - 0.70 pmhos chloride <50 ppb sodium (Na) <10 ppb 2 ~2 1
0 In order to provide assurance that feedwater purity was maintained on AVT, condenser integrity was continually monitored by eddy current examination and preventative plugging of'damaged tubes during shutdowns. As a further measure, full flow, deep bed condensate polishers were installed and put into operation in January 1978. Blowdown rates were increased to 70 gallons per minute per steam generator in February 1979.
The following blowdown chemistry comparison demonstrates the continued improvement in bulk water chemistry over the last eight years:
Parameter 1974-1977 1978 1981 Cation Cond., pmhos 0.7-2.5 0.2-0.4 0.12-0.2 Chloride, ppb 50 < 10 3-5 Sodium, ppb 5-15 5-15 3-8 Silica, ppb 20-50 15-30 5>>10 pH 8.6-9.0 8.7-8.9 8.7-8.9 Since November 1969, oxygen concentration has generally been less than 5 ppb. There have been several times over the last twelve years when dissolved oxygen has been as high as 40 ppb for several days. From July 1978 to the present, feedwater dissolved oxygen has been less than one (1) ppb.
Operating chemistry has been closely monitored throughout plant life and has progressively improved over the operating history of Ginna Station. Participation in the Steam Generator Owners Group with the Electric Power Research Institute has provided valuable information regarding operating chemistry. Presently, Ginna Station's secondary water chemistry and feedwater purity are well within the Owners'roup Guidelines.
2.2.2 ~Plu ~in Ginna Station's first pluggable defects were identified in March 1974 by an eddy current examination of both steam generators. The pluggable defects were found only in the A-Steam Generator and were confirmed to be wastage of the Inconel 600 tubing. Within three years after conversion to the 'AVT water chemistry, the plugging of tubes in the A-Steam Generator for secondary side attack was essentially eliminated. See Table 2.1 and Fig. 2.3 for further definition of the A-Steam Generator tube plugging history.
The B-Steam Generator has had a different plugging history. Its first pluggable defect did not occur until after the AVT conversion.
The first pluggable defects were identified in March 1975 and were characterized as stress corrosion cracking (SCC). SCC occurred because of the molar ratio shift (Na/PO ) resulting in the sludge after conversion from phosphate contrSl to the all volatile treatment (AVT) chemistry control. This shift resulted in a caustic environment being formed in previously phosphated 2%2 2
sludge since, on AVT, there is no addition of phosphate to buffer the sodium hydroxide. As with the A-Steam Generator, within a few years after AVT conversion, secondary side corrosion indications subsided. However, beginning in 1976, defects were found in the periphery area. These defects were mainly in the wedge areas (areas where the support plates are attached to the shell of the steam generator) within the first four (4) inches above the tube sheet. However, four (4) tubes adjacent to the Number 4 wedge area were found to have indications about 24 inches above the tube sheet. The majority of these defects were interpreted as internal diameter indications based on the eddy current signals.
In the spring of 1979, the first indications of tubesheet crevice intergranular attack (IGA) were identified in the B-Steam Generator.
Since that time, a total of 74 indications of crevice IGA have been identified, 16 of which have been sleeved and the others plugged. Table 2.2 and Figure 2.4 provide further definition of the B-Steam Generator tube plugging history. Table 2.3 is a detailed listing of the periphery defects including tubes that have been plugged one or two rows in from the periphery.
2.2.3 Secondar Side Modifications In March 1975, modifications were made to the steam generators to improve the lateral flow velocities across the tube sheet and to improve blowdown efficiency. The purpose of increased flow velocity and blowdown was to reduce the sludge pile area and volume on the tube sheet, thereby reducing the concentrating medium where corrosion of the tubes had been seen.
The modifications consisted of the following:
a) Modification of the feedring so that 80% of the feedwater would flow down the hot leg side of the steam generator and the remaining 20% would flow down the cold leg side.
b) Removal of the downcomer flow resistance plate to increase the recirculation ratio from 2.36 to 4.73.
c) Replacement of the moisture separator swirl vane orifice rings with smaller diameter rings to increase the moisture separator efficiency.
d) Installation of blowdown lane flow blockers to prohibit flow from bypassing the tube bundle and coming into the blowdown lane directly from the wrapper annulus.
Due to the increased recirculation ratio, the total moisture carryover was higher than desired. Therefore, a modification was performed in February 1976 consisting of the installation of perforated plates on the moisture separator demister end plates.
2 ~2 3
To provide further assurance that water hammer due to feedring drainage would not occur, another modification of the feedrings was performed in February 1979. This modification involved plug-welding all the bottom drain holes on the feedring and installing J tubes on top of the feedring. The J tubes were located so as to maintain the 80% to 20% feedwater flow split.
between the hot and cold legs, respectively.
During the March 1982 refueling and maintenance outage, further moisture separation equipment modifications have been implemented to increase their effectiveness and efficiency. These modifications involved installation of upper deck plate relief with directional chimneys and steam flow redirection pagodas over the swirl vanes, and the addition of larger demister drains.
2.2-4
TABLE 2.1 GINNA STATION A-STEAM GENERATOR TUBE PLUGGING HISTORY DATE WASTAGE SCC* PITTING MANUF. TOTAL In Factory March, 1974 19 19 November, 1974 March, 1975*** 46** 46 February, 1976 39 39 April, 1977 13 13 April, 1978 April, 1980 TOTALS 73 122
- - SCC (caustic stress corrosion cracking.)
- one tube leak of less than 0.1 gallons per minute
- - plugging criteria changed from 50% to 40% wall penetration 2.2-5
TABLE 2.2 GINNA STATION B-STEAM GENERATOR TUBE PLUGGING HISTORY DATE WASTAGE SCC+ IGA++ PERIPHERY TOTAL March, 1975+++
January, 1976 (2)* (2)
February, 1976 April, 1976 15* 15 April, 1977 July, 1977 January, 1978 April, 1978 15 15 February, 1979 December, 1979 13 April, 1980 November, 1980 3**
May, 1981 15 January, 1982 13 16 TOTALS 15 74** 49 147 (2) " these 2 tubes were wastage indications below the top of the tube sheet on the periphery 5 tube leaks less than 0.1 gallons per minute
- - 28 tubes of 74 were not above plugging limit of 40% wall penetration
- - R42 C55 tube rupture
+ - SCC (caustic stress corrosion cracking)
++ - IGA (tube sheet crevice intergranular attack of the tubes)
+++ - plugging criteria changed from 50% to 40% wall penetration 2.2-6
TABLE 2.3 GINNA STATION B-STEAM GENERATOR PERIPHERY AREA DEFECTS ORIGINAL EDDY CURRENT INTERPRETATION January 76. R39 C69 O.D.'s top of tube sheet (TTS)
R40 C68* I.D.'s 2" below TTS May 76 R8 C92 I.D. at TTS R9 C91 I.D. 3" above TTS R10 C91 above R11 C91 I.D. at TTS R12 C91* I.D. at TTS R13 C90 Distorted tube sheet entry R14 C90 I.D. 2"-4" above TTS R15 C90 I.D. 2" above TTS R15 C89 OK (No Defect)
R16 C89 I.D. 2" above TTS R17 C89 I D 2" above TTS R30 C15 47% 0 D 2ii above TTS R31 C15 43% O.D. 3" above TTS R32 C15 I.D. at TTS R33 C15 I.D. at TTS Jury 77 R45 C54* Bulge and I.D. signal 2"-4" above TTS R44 C55 Bulge and I.D. signal 2"-4" above TTS R44 C56 Bulge'and I.D. signal 2"-4" above TTS R44 C57 Bulge and I.D. signal 2"-4" above TTS R44 C58 Bulge 2"-4" above TTS R12 C2 O.D. above TTS January 78 R45 C53 Bulge and I.D. 2"-4" above TTS R44 C54* 100% Defect (July 77 I.D. above TTS)
April 78 R45 C50 I.D. at TTS R45 C51 I.D. at TTS R45 C52 Bulge,2"-4" above TTS, Defect below TTS R44 C52 Damaged when R45 C52 was pulled R44 C53 Distorted TTS R43 C58 I.D. at TTS R43 C59 I.D. at TTS R43 C60 Noisy I.D. signal at TTS C61 Noisy I.D. signal at TTS C66 Many I.D. signals at TTS R40 C67 I.D. at TTS R39 C68 O.D. and I.D. above TTS R39 C70 I.D. at TTS R38 C71 I.D. at TTS R38 C72 I.D. at TTS 2 2 7
TABLE 2.3 (continued)
February 79 R43 C55 Distorted tube sheet entry R43 C56 O.D. and I.D. above TTS R35 C75 O.D. at 42 Support plate (cold leg)
'3%
R28 C12 47% O.D. at 41 Support plate (cold leg)
December 79 R43 C54* 100% Defect R43 C57 80% O.D. above TTS April 80 R45 C48 46% O.D. 24>> above TTS R45 C49 36% O.D. 24>> above TTS R32 C16 above April 81 R43 C53 80% O.D. above TTS February 82 R42 C55 Burst R45 C46 41% O.D. 24>> above TTS
'45 C47 49'~ O D 24<< above
- - leaks less than 0.1 gallons per minute 2.2-8
/TEAM OUTLET TO TURBINE GENERATOR DEMISTERS SECONDARY MOISTURE SEPARATOR SECONDARY MANWAY r g , ~ "
ORFICE RINGS UPPER SHELL SWIRL VANE PRIMARY MOISTURE SEPARATOR FEEDWATER.RING FEEDWATER INLET ANTIVIBRATION BARS TUBE BUNDLE Ij'~, PPgg fl!1I)
LOWER SHELL ll!i!
WRAPPER
))jtI %BE SUPPORT PUITES BLOWDOWN UNE SECONDARY HANDHOLE TUBE SHEET PRIMARY MANWAY TUBE tANE BLOCK PRIMARY COOtANT OUTLET PRIMARY COOLANT INLET QK SERIES 44 STEAM GENERATOR NOT TO SCALE 0$ 9 FIGURE 2.1 2.'2-9
ell
- .v>> ~
'GEE ~
Steam Generator HOT LEO (INLET)
&878i838191115757I L9LTL515LI5957 55 535I49474543 4l 393735333I 292725232I I9 I7 l5 l3 II 9 7 5 3 I CQLUMNS
'92 9o888L 8 5 52 412 323 l816 l4 l2 IO S6 WEDGE AREA WEDGE AREA.
NUMBER 4 NUMBER 3 45 42 43 4I 40 WEDGE AREA 38 NUMBER WEDGE AREA 36 NUMBER 2 33 32 3I 29 28 26 25 24 22 23 2l 20 I
l7 INEDGE AREA'R' I5 IIUMBER 1 I
'EDGE AREA IO II NUMBER e 8 7
5
.RGWS (oW T)IS A MIRROR IMARS INITR ~
~MA ylAY FIGURE 2.2
ASIA e RGSE Steam Generator A IHLKT PLUOOKO TUKK MAP oil N8785&61911757'57I L9L'7L5L3II59575553514947454541 393735333129272523211917 15 1311 9 7 5 3 I COLUMNS 9'2 90 8Z 74 7Z 8 LLL4 L2 0 54 525 4L 442, 4 3 323 Z8 22 18 16 1412 10 8 6
~ ~ ~
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30 28 26 24 22 20 18 16 14 12 I
10
~MANWAY . ~ PAKVIOUKLY PLUOOKD TUbdd NOZZLE ~ ROWS'IO.
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8 Sl RGSE Steam Gener r 8 lNLET PLUOOEO TU8E 'MAP 9l NS7K %679117S7571L94745LEl 15957 55 535149474543 4l 3937353331 29272523211917 15 1311 9 7 5 3 I CQl UMNS 82 0 SLl 4L2 0 54 52 442 323 2 1816 14 1210 4 2 44
~ ~ ~ ~ ~ ~ ~ 42
~ ~ ~ ~
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32 30 28 26 24 22 20 I
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10 8
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~ MANWAY ~ PLuGGED TUB~S NOZZLE ~ ROWS X SLEEVED TUBES
'IG. 2.4
3.0 STEAM GENERATOR INSPECTION RESULTS 3.1 Edd Current Examination The inspection of the B-Steam Generator after the January 25th tube rupture started with a hydrostatic leak test to identify the burst tube. By visual inspection during the leak test, tube R42 C55 in the B hot leg (inlet) was identified as the leaking tube. No other tubes were found to be leaking from this leak test. After visual verification, a single frequency 400 kHz eddy current examination was made of the leaking tube by hand probing from the primary manway. A large volume defect was found approximately 5 inches in length running from 3 to 8 inches above the tube sheet.
The method used for the multifreguency eddy current, examination of the total steam generator was the same method that has been used for the last four inspections. This state-of-the-art method includes techniques that provide maximum sensitivity to the types of damage which are being identified in steam generators throughout the industry. Specifically, both differential and absolute technicpes are used as follows:
a) 400 kHz differential b) 200 kHz differential c) 210 kHz absolute d) 100 kHz absolute In order to increase the sensitivity in areas where support plates and tube sheet indications normally mask actual tubing defects, mixing of both the differential and absolute frequencies respectively, was utilized.
The scope of the multifreguency eddy current examination included 100% of both steam generators'ot leg (inlet) tubes, all periphery tubes, and a sampling of the cold leg (outlet) tubes of both steam generators. A tube examination consists of examining a tube from at least the first support plate through the tube sheet. A sampling of tubes over the U-bends and to the sixth support plate level was also examined.
Results of the multifrequency eddy current examination of the A-Steam Generator hot and cold legs (inlet and outlet) and the B-Steam Generator cold leg (outlet) revealed no pluggable indi-cations or any noticeable changes from previous inspections.
There were no pluggable indications or any noticeable changes from previous inspection results above the first support plate in either steam generator. However, the B-Steam Generator hot leg (inlet) did reveal that further degradation had occurred in several tubes since the last inspection. There were no indications found similar to the burst tube R42 C55. Two tubes, R45 C46 and
'45 C47, had indications 24 inches above the tube sheet of 41%
and 49% through the wall, respectively. These indications were approximately mid-span between the tube sheet and the first 3.1-1
e support plate. Thirteen tubes were found with defects in the crevice, similar to the defects from crevice IGA which have been identified in the past. The following is a list of those tubes:
- 1. R9 C44 - 14 inches below the top of tube sheet (TTS)
. 2. R16 C42 12 to 14 inches below TTS
- 3. R18 C39 4 to 12 inches below TTS
- 4. R20 C45 - 10 inches below TTS
- 5. R20 C44 - 10 inches below TTS
- 6. R21 C56 - 10 inches below TTS
- 7. R21 C43 - 4-12 inches below TTS
- 8. R24 C56 4 inches below TTS
- 9. R24 C48 - 6-10 inches below TTS
- 10. R33 C59 - 17 inches below TTS to rolled transition ll.
12.
R35 R35 C54 C53 16 inches below TTS 'to 12 inches below TTS to rolled rolled transition transition
- 13. R38 C40 - 16 inches below TTS to rolled transition Five of these crevice indications have not changed since the Spring 1981 inspection.
3~1 2
3.2 Prof ilometr Examination Assessment of any geometric conditions that may have been associated with the burst tube, or other tubes in the vicinity of the burst tube, was accomplished by use of the Babcock and Wilcox'ompany's profilometry equipment. This procedure utilizes strain gauges with fingers attached which are calibrated to known deflections of larger or smaller diameters than the nominal inside diameter of the tubes.
Approximately 75 tubes were examined including the burst tube (R42 C55), tubes around the burst. tube, and a sampling of tubes on the periphery.
The results of, this inspection revealed about 20 tubes that had a dent at the first support plate. These dents were generally less than 10 mils. In addition, the burst tube had bulged in the area of fracture as had been expected. These results did not provide any insight into the cause of the burst.
3~2 1
3.3 Fiber 0 ties Ins ection As part of the inspection program, fiber optics inspections were performed from the primary and the secondary side of the burst tube. The secondary side inspection included probing the columns on either side of the burst tube, and then coming around the periphery of the steam generator on the tube sheet and viewing into the bundle in the area of the burst tube.
The primary side probing of the burst tube revealed a diamond shaped burst approximately 4 inches long and 0.75 inches wide at its widest point. Secondary side inspection results from the column probing revealed no abnormal observations.
As the inspection was progressing around the periphery, abnormal damage was observed to previously plugged tubes around rows 44 and 45, columns 58 to 53. Also, as the fiber optics equipment
. was being removed from the steam generator, a foreign object was identified around the area of R25 C85. The foreign object identified was a carbon steel plate 0.5 inches thick by 4.18 inches wide by 6.31 inches long. This plate was removed along with another piece of carbon steel approximately 0.050 inches thick by 0.6 inches wide by 4 inches long. Because of these results, further inspection was deemed necessary, utilizinq equipment and procedures that would allow for a greater field of vision with higher resolution of the steam generator periphery tubes.
3 ~ 3't 1
3.4 Television Video Ins ection After the fiber optics inspection results. became available on February 10 and 11, equipment was located capable of performing a video inspection within the geometric constraints of the steam generator periphery. A Westinghouse Underwater Reactor Vessel Inspection Video System was selected for the inspection. After three days of mock-up work and procedure development, the inspection technique was implemented in the B-Steam Generator on February 14.
The video inspection consisted of, first, scanning the peripheiy tubes around the total circumference of the steam generator hot and cold legs and then, second, scanning the columns perpendicular to the-tube sheet blowdown lane. These two areas were inspected in both the A and B-Steam Generators, as defined by Fig. 3.4-1.
Scanning the columns perpendicular to the tube sheet blowdown lane did not reveal any abnormalities in either steam generator.
Foreign objects were found in both steam generators. Tube damage was only found in the B-Steam Generator hot leg. The tube damage identified was mainly associated with previously plugged tubes in the Number 4 and 6 wedge areas. In addition, some minor scrapes and dings were identified on tubes located near R40 C68. No defects were found on tubes in the Number 2 wedge area. (This was confirmed by subsequent, more detailed, fiber optics examination.)
Complete inspection results for the B-Steam Generator hot leg are documented in Table 3.4-1.
3.4-1
Page 1 of 3 TABLE 3.4-1
.DATA REVIEW B-STEAM GENERATOR HOT LEG Tube Results Location
~see Fag. 3.4-2)
R1C92 OK R2C92 OK R3C92 OK R4C92 OK R5C92 OK R6C92 OK R7C92 OK R8C92 COLLAPSED WITH HOLE PLUGGED A surface) -
-(Ripples R9C91 O.D. DAMAGE on PLUGGED B R10C91 O.D. DAMAGE WIPED AREA - SCRAPED (Ripples)
PLUGGED R11C91 COLLAPSED WITH HOLE (Ripped) 4-5" above T.S. - PLUGGED D R12C91 COLLAPSED WITH HOLE 4" Above T.S. - PLUGGED E R13C90 -
O.D. DAMAGE PLUGGED F R14C90 COLLAPSED WITH O.D. DAMAGE - DEPRESSION PLUGGED R15C90 O.D. DAMAGE WITH HOLE (Collapsed 5 Ripples)
PLUGGED H R16C89 MINOR O.D. DAMAGE - PLUGGED I R17C89 -
O.D. DAMAGE PLUGGED J R18C88 OK R19C88 OK R20C88 OK R21C87 OK R22C86 OK R23C86 OK R24C85 OK R25C85 OK R26C84 OK R27C83 OK R28C82 OK R29C82 OK R30C81 OK - SMALL SHREAD OF TUBING ON TUBE SHEET K STAY ROD OK R33C78 OK I R34C77 OK R34C76 OK R35C76 OK R35C75 OK - PLUGGED R35C74 OK R36C74 OK R36C73 OK R37C73 OK R37C72 OK R38C72 OK - MINOR O.D. DAMAGE - PLUGGED L R38C71 OK - MINOR O.D. DAMAGE - PLUGGED M R38C70 OK 3.4-2
t Page 2 of 3 DATA REVIEW B-STEAM GENERATOR HOT LEG Results Location
~see Fag. 3.4-2)
R39C70 OK - -
O.D. DAMAGE PLUGGED N R39C69 OK - O.D. DAMAGE PLUGGED 0 R39C68 OK - -
O.D. DAMAGE PLUGGED P R40C68 OK - PLUGGED Q R40C67 OK - PLUGGED R R41C66 OK - PLUGGED S R41C65 OK O.D. DAMAGE - DINGS T R41C64 OK -
O.D. DAMAGE . DINGS U R42C64 OK - -
O.D. DAMAGE DINGS V R42C63 OK - -
O.D. DAMAGE DINGS W R42C62 O.D. - -
O.D. DAMAGE DINGS X R43C61 SCRAPED ON O.D. >> PLUGGED Y R43C60 SCRAPED ON O.D. - PLUGGED 2 R43C59 -
O.D. DAMAGE PLUGGED AA R44C58 COLLAPSED, O.D. DAMAGE - PLUGGED AB R44C57 COLLAPSED, O.D. DAMAGE - PLUGGED AC R44C56 COLLAPSED, MISSING - PLUGGED AD R44C55 COLLAPSED, O.D. DAMAGE - PLUGGED AE R45C54 COLLAPSED, MISSING - PLUGGED AF R45C53 COLLAPSED, O.D. DAMAGE - PLUGGED AG R45C52 PULLED TUBE AH R45C51 OK MOTTLED (RIPPLED O.D.) - PLUGGED AI R45C50 OK - PLUGGED R45C49. OK - PLUGGED R45C48 OK - PLUGGED R45C47 OK - PLUGGED R45C46 OK -. PLUGGED R45C45 OK R45C44 OK R45C43 OK R45C42 OK R45C41 OK R45C40 OK WIRED'ND R45C39 OK R44C38 OK R44C37 OK R44C36 OK R44C35 OK R43C34 OK PIECE OF PLATE 3 5 X 1 5 X 5 AJ R43C33 OK R43C32 OK R42C31. OK R42C30 OK
- 3. 4-3
Page 3 of 3 DATA REVIEW B-STEAM GENERATOR HOT LEG Results Location
~see Fag. 3.4-2)
R42C29 OK R41C28 OK R41C27 OK R40C26 OK R40C25 OK R39C24 OK R39C23 OK R38C21 OK R37C20 OK R36C19 OK R35C18 OK R35C17 OK R34C16 OK R33C16 OK I R33C15 OK - PLUGGED, SMALL SHREAD OF TUBING BETWEEN R32 C15 Sc STAY BAR STAY BAR OK R30C13 OK R30C12 OK R29Cll OK R28C11 OK R27C10 OK R26C9 OK R25C8 OK R24C8 OK R23C7 OK R22C7 OK R21C6 OK R20C5 OK R19C5 OK R18C5 OK R17C4 OK R16C4 OK R15C3 OK R14C3 OK R13C3 OK R12C2 OK - PLUGGED AL R11C2 OK R10C2 OK R9C2 OK R8cl OK R7C1 OK R6C1 OK R5C1 OK R4C1 OK
'R3C1 OK R2C1 OK Rlcl OK 3.4-4
FIGURE 3.4-1 Steam Generators A and B Secondary Side Tube Sheet Periphery Video Ins ection Pro ram Scan periphery along tube sheet of Total circumference of Steam Generator (includes hot and cold legs).
- 2. Scan columns perpendicular to tube sheet blowdown lane.
I5 It II 0 4I 5 II 5I II 5I 5 5 0 4 aII 4C kCI Ml A.S aII a54 a II
- ~ att e ' +/
/
/ 455 4,II
- 55 I Ml l It 4 IC
- II
+p// A4 p A4
+5 Op AC 45
~4 aI SI I
4'+ 4.IaS t<<<55o
<<155
%b r
B-Steam Generator Rl 8}818}856191115&7II9L1LSt~3L}5'}5155535}494145434}393735333I 292725232II917 I5 I3II 9 7 5 3 I COLUMNS
'929 888 82 S l 74Zo LI. 4 54 52 4L 4 IZ 3 3L 0323 I8 I I4 I2 IO
- RG6E B-STEAM GENERATOR HOT LEG
~ ~ I ~
44 45 42 40 38 36 4
32 33 3I 30 28 27
~ ~
25 23 22 S
20 Ig 4 g 17 E I6 D I3 5 I2 IO 8
6 4
2 ROWS
~MANWAY NOZZLE~
FIGURE 3.4-2
3.5 Forei Ob 'ects 3.5.1 A-Steam Generator No foreign objects were found in the A-Steam Generator hot -leg.
The following foreign objects vere found in the A-Steam Generator cold leg.
a) A piece of wire 0.0375 inches in diameter, 11.125 inches long, non-magnetic, stainless steel located near R34 C77.
b) A piece of wire 0.1265 inches in diameter 4.5625 inches long, magnetic, carbon steel weld rod, located near R44 C38.
c) Piece of metal with portion of weld. Irregular shape approxi-mately 0.5 inches thick by 0.75 inches vide by 1 inch long.
Magnetic material, carbon steel plate segment. Located near R44 C38.
3.5.2 B-Steam Generator The following is a listing of foreign objects found in the B-Steam Generator. Figure 3.4-2 illustrates the primary locations where the foreign objects were found in the B-Steam Generator hot leg.
a) Piece of magnetic carbon steel plate 0.5 inches thick by 4.18 inches wide by 6.31 inches long. Located near R25,C85 area.
b) Piece of magnetic carbon steel plate 0.5 inches thick by 1.5 inches wide by 3.5 inches long. Initially located near R45 C46 area.
c) Piece of magnetic carbon steel plate oval shape 0.5 inches thick, minor axis 2.0 inches with major axis 2.375 inches.
Located wedged beween R45 C53 and R44 C53.
d) Piece of magnetic carbon steel strip 0.050 inches thick by 0.6 inches wide by 4 inches long. Located near R25 C85 area.
e) Piece of copper tubing, approximately 0.25 inches in diameter by 1.062 inches. long. Located near R45 C47 area.
Piece of welding electrode 0.18 inches in diameter by 2 inches long. Located near R43 C34 area.
g) Four pieces of welding slag small ball shapes less than 0.5 inches in diameter.
I h) Tvo pieces of material small ball shapes less than 0.25 inches in diameter.
3.5-1
i) Small pieces of Inconel tubing were also damaged tubes of various lengths.
identified from These were located in the
~
Number 4 wedge area. Two other pieces were identified, one near R30 CSl the other near R33 C15.
- 3. 5-.2
- 4. 0 STEAM GENERATOR TUBE FAILURE ANALYSIS PROGRAM 4.1 Introduction 4.1.1 General An extensive failure analysis program has been developed to determine the mechanisms which resulted in the rupture. of tube R42 C55 on January 25, 1982. The purpose of this program is to provide analytical and experimental information relative to the role that, various, postulated failure mechanisms had in the tube rupture. This information is then used to establish the most probable failure mechanisms, and take the corrective actions necessary to preclude recurrence.
4.1. 2 Procrram The failure analysis program was developed based on the operating history of the B-Steam Generator, inspection results, and previous analysxs and test results. The program was designed to assess the various mechanical mechanisms which may have been involved in the tube failures. These include buckling, collapse, fatigue, severing, wear, and burst. The program includes assessment of the effect that foreign objects may have had in the failure mechanism.
The program consists of the following elements:
a) postulated failure mechanism b) data review c) metallurgical examinations d) analyses e) model testing f) laboratory testing The majority of the analysis and testing work has been performed by Westinghouse. The collapse and fatigue testing is being performed at the Research and Development Laboratories in Churchill, Pennsylvania. The flow model testing was performed at the Westinghouse Engineering Test Facility in Tampa, Florida. Sections 4.5 through 4.8 of this report contain a detailed description of the analyses and test's'performed by Westinghouse together with the associated results and conclusions from that work.
4.1-1
In addition to Westinghouse, analysis and testing has been performed by Combustion Engineering. The analysis work was performed under contract to the Electric Power Research Institute, Inc. (EPRI) and the Steam Generator Owners Group. Appendix B to this report contains a summary of that work. Testing relative to the effects of axial load and loose part impacting on a steam generator tube's propensity for local buckling is presently in progress at Combustion Engineering s Chattanooga, Tennessee facilities. A description of the test equipment, parameters, and methodology is included in Appendix C to this report. The results of this work will be provided as an Addenda when they are available.
4.1 2
4.2 Postulated Failure Mechanisms 4.2.1 Initial Plu in As stated in Section 2.2.2 of this report, plugging of tubes in the B-Steam Generator periphery began in 1976. The tubes were plugged for a combination of internal and outside diameter eddy current indications. The indications occurred between the tube sheet and first support plate. Since many of the indications were internal diameter, and these tubes were in a region of the steam generator where very little sludge can accumulate, corrosion is not postulated to be a mechanism involved in the failure.
Therefore, the most probable cause of the initial tube plugging is postulated to be some adverse lateral or axial mechanical loading mechanism on the tubes.
1 4.2.2 ~Colla se Video inspection and subsequent tube removals have shown the presence of collapse areas on several plugged tubes on the B-Steam Generator hot leg periphery. Based on the location of these tubes in the outermost tube rows, and the time at which they were originally plugged, it appears that collapse was the next step in tube degradation following the initial plugging. The factors which may have contributed to the collapse are external pressure, lateral impact, and axial load. Lateral impact, of the tube could have resulted in ovalization, wall thinning, or cyclic stresses.
Axial load could have resulted in additional stresses in the tube. Since the tubes were previously plugged, they were under external pressure during normal plant operation. Ovalization, wall thinning, and/or mechanical stress could have weakened the tube to the point where pressures.
it collapsed under normal operating 4.2.3 Severance /
Several tubes in the Number 4 wedge area of the B-Steam Generator were found to be severed. This included tubes both with and without evidence of collapse. The two outermost tubes which had collapse areas were severed at the tube sheet and first support plate. The remaining tubes were located primarily in inner rows and had been severed at the tube sheet only. Therefore, appears that severing was the next step in the failure process it following collapse. Severing of the tube is postulated to have occurred as a result of either wear or fatigue loadings from flow induced vibration and/or lateral impact. The lateral impact could have been either from another severed tube or a foreign object. Flow induced vibration could have resulted from weakening of the tube due to a collapse area, and/or lowering of natural frequency as a result of the collapse and/or axial load.
- 4. 2-1
4.2.4 Wear Many tubes, particularly in the Number 4 wedge area, were found to have outside diameter (O.D.) wear. The wear appears to be primarily the result of one tube rubbing on another. This is postulated to have occurred from the severed tubes. In a few cases, very localized wear areas were seen. These appear to have been caused by a foreign object. Wear on a previously plugged tube could have weakened it to the point where it severed and subsequently caused impact or wear on additional tubes. Wear on an active tube would result in either a defect which was detected
.by eddy current and the tube plugged; or enough wall thinning over a sufficient area to cause the tube to burst from internal pressure.
- 4. 2. 5 ~5ummar Figure 4.2-1 illustrates the various failure mechanisms, and their sequence, which are postulated to have resulted in the rupture of R42 C55. A foreign object, axial loads, and pressure are postulated to have combined in leading to. a sequence of events which led to plugging of many tubes on the B-Steam Generator periphery, and eventually to the rupture of one of them. The intermediate steps leading to the rupture appear to have-included collapse, severing, and wear. Most of this intermediate tube degradation occurred subsequent to plugging and, therefore, could not be detected by eddy current examination. The analyses and tests which are described in the following sections have been performed to substantiate the postulated failure mechanisms described here.
- 4. 2-2
0 0
Mechanical Load Active Tube Tube Plugged Collapse Vibration Shredding Sever Wear Wear Plugged Tube Active Tube Plugging Burst Figure 4.2-1 Postulated Failure Mechanism Sequence 4.2-3
~ ~
/ I ~ ' ~ ~ ~ c ~ ' ~
4.3 Data Review 4.3. 1 Puzu>ose The purpose of this section is to review the examination and repair history of the peripheral tubes in the B-Steam Generator.
The results of this review are then related to the results of the February 1982 examinations. Only one peripheral tube in the A-Steam Generator has been plugged, and that tube was plugged during steam generator fabrication. No peripheral tubes have been plugged in the A-Steam Generator during operation.
4.3.2 Examination and Re air Histo Ginna Station began operation in late 1969. Steam generator eddy current inspections were performed in April 1972, March 1974, November 1974, and March 1975. There were no indications of any tube degradation in any peripheral tubes found during any of these inspections.
In January 1976, the plant was shut down with a periphery tube leak in the B-Steam Generator. The leakage, through tube R40 C68, was less than 0.1 gpm. R40 C68 was determined to have I.D.-type indications 2 inches below the top of the tube sheet and was plugged. R39 C69 had O.D. indications at the top of the tube sheet and was also plugged.
In February 1976, the plant was shutdown for refueling, maintenance, and steam generator inspection. There were no eddy current indications in the B-Steam Generator peripheral tubes.
In April 1976, shortly after startup from the February 1976 outage, a tube leak occurred in the B-Steam Generator. The leak rate was less than 0.1 gpm from the tube R12 C91, near the tube lane on the manway side of the bundle in the Number 6 wedge area. Eddy current inspection revealed I.D.-type indications at the top of the tube sheet. Neighboring tubes on the periphery in this wedge area were shown by eddy current inspection to contain signals which were characterized also as I.D.-type. These also were located at, or just above the top of, the tube sheet. In addition, a few tubes in the Number 6 wedge area also produced eddy current signals at the first tube support plate. These were characterized as dents of various sizes, all less than 10 mils.
All peripheral tubes with any I.D.-type indications in this area is found p fthhtlg'5l'th-in Figure 2.4. Wedge area locations are defined tin Figure 2.2.
4.3-1
were plugged during this outage. (Also plugged at this time were 4 tubes on the opposite side of the bundle, in Column 15, Rows 30-33, which is in the Number 2 wedge area. Two of these, R30 C15 and R31 C15, exhibited O.D. indications, 45% through-wall, about 2 inches above the top of the tube sheet. The other 2 tubes displayed I.D.-type of signals at the top of the tube sheet.)
Eddy current inspection of selected tubes in the Number 4 wedge area (90'o the tube lane and furthest from it, at the periphery) showed no significant indications during the April 1976 outage.
During the April 1976 inspection, an individual reached into the secondary side of the B-Steam Generator through the 6 inch blowdown lane handhole to probe the Number 6 wedge area. No evidence of external distress was felt on the peripheral tubes and no unusual conditions or foreign objects were noted between the tubes and the steam generator shell.
The plant then operated until April 1977. In April 1977, inspections of the Number 4 wedge area revealed one tube, R45 C54, to have two I.D.-type indications, one slightly above and one slightly below the top of the tube sheet. Three neighboring tubes did not exhibit recordable indications at this investigation. The unit was returned to service in mid-May 1977.
Shortly thereafter, on July 5, 1977, a minor leak (0.09 gpm) occurred in the B-Steam Generator. The leaking tube was identified as R45 C54, which had been identified as having indications the preceeding April. In the July inspection, R45 C54 exhibited bulge and I.D.-type indications. Inspection of neighboring tubes in the Number 4 wedge area indicated that four tubes in Row 44 (Columns 55-58) also contained bulge type indications and I.D.-type signals in the tube sheet area. These five tubes plus the tube at R12 C2, which had an O.D. indication above the tube sheet, were plugged and the unit was returned to service.
In January 1978, a leak occurred in the Number 4 wedge area at tube R44 C54, adjacent to the columns that were plugged the preceding July. Following the plugging of this tube and the adjacent tube (R44 C53), which displayed a bulge signal and an I.D.-type indication above the tube sheet, the unit was returned to service and operated until the refueling and maintenance outage of April 1978.
In April 1978, several tubes in the Number 4 wedge area were identified by eddy current testing to exhibit I.D.-type signals near the top of the tube sheet. Several dent-type indications at the first support plate were identified on some (but not all) tubes containing I.D.-type signals. The dents were all less than 10 mils. Several additional peripheral tubes in an area centered at R40 C69 (about 3/4 of the peripheral distance from the Number 6 wedge area to the Number 4 wedge area) also exhibited the I.D.-type signals.
4.3-2
0 During the April 1978 outage, one of the Number 4 wedge area tubes, R45 C52, was removed for laboratory investigation.
Metallurgical, microanalytical, and metallographic studies. were performed and are reported in an Electric Power Research Institute Report (Ref. 4.3-1). No evidence of I.D.-type degradation (such as primary water stress-corrosion cracking, other corrosion .
phenomena, or I.D. mechanical processes) was detected in this analysis. The tube exhibited a 15-mil diametral "bulge" above the top of the tube sheet in an area containing many small peen-like marks and "ripples" on the outside diameter (O.D.) surface. The rippled area also displayed a slight amount of wall thinning (approximately 15% maximum) and a slight increase in micro-hardness.
The features that were observed on the O.D. surfaces were mechanical appearing in nature and, except for the "bulge", are of the type which could have been caused by a "peening-like" action of a solid object interacting with the O.D. surface. Information available to the analysts was that the O.D. mechanical features faced the tube lane and not the periphery of the bundle. (It is conceivable, however, that, the orientation of the investigated tube sections may not have been unambiguously established.)
Following plugging in April 1978 of all I.D.-type indications in the Number 4 wedge area and in the R40 C69 area (15 total tubes),
the plant returned to service.
Inspections in February and December 1979 resulted in plugging several more tubes in the Number 4 wedge area. The tube at R43 C56 (February 1979) and the tube at R43 C57 (December 1979) exhibited large O.D.-type indications 2-3 inches above the top of the tube sheet. In April 1980 and May 1981, three Number 4 wedge area tubes were plugged for O.D. indications. The 1980 tubes, R45 C48 and R45 C49, exhibited approximately 40/ O.D. indications about 24 inches above the top of the tube sheet.
Following the January 1982 event in the B-Steam Generator, the original eddy current findings from the Number 6 and Number 4 wedge areas plugging were reviewed and reevaluated. The peripheral tubes were also characterized by a video tape examination conducted in February 1982. Tables 4.3-1 and 4.3-2 present the results of the recent reevaluation for the Number 6 wedge area and the Number 4 wedge area, respectively. Table 4.3-2 also presents eddy current interpretations from inspections prior to the plugging date for the Number 4 wedge area. (Tubes in the Number 6 wedge area yielded no eddy current indications at inspections prior to the outage at which they were plugged). Also provided in both tables is a summary of the February 1982 video and visual inspections. Figure 4.3-1 is a schematic representation of the plugging history in the Number 4 wedge area.
4.3-3
4.3.3 Evaluation Many of the plugged tubes in both the Number 4 and Number 6 wedge areas were observed to be externally damaged during the February 1982 videoscan inspections of the bundle periphery. Eddy current data was interpreted at the time of plugging, and have been confirmed through reevaluation now, as being a bulge or I.D.-type indication.
It is concluded that the defects which caused the original tube plugging in the Number 4 and Number 6 wedge areas cannot be established from a review of the eddy current inspection data alone. The I.D.-type of eddy current indications that were observed in the 1976 inspection of the Number 6 wedge area were of appreciable amplitude. Such signals could be attributable to:
(a) Actual I.D. discontinuities such as cracks (axial or circumferential) pits, thinning, magnetite deposition (on the I.D.), or permeability variations.
(b) A localized bulge or diametral increase.
(c) The presence on the O.D. of a highly conducting layer, such as copper.
Examination of the Number 6 wedge area tubes is currently in progress using metallurgical failure analysis techniques, including radiography and metallographic sectioning. Radiography of tubes R14 C90, R15 C90 and R12 C91 (the 1976 leaker) has not indicated any cracks which could have accounted for the original I.D.-type signals. Sectioning of tube R15 C90 over a length from 1 to 6 inches above the top of the tube sheet did not .reveal any I.D.
discontinuities which could have generated the 1976 eddy current signals.
The observations of O.D. indications of hot leg peripheral tubes in the Number 4 wedge area was made after the I.D.-type of indications were first observed in the same wedge area. The first observation of indications in this area was in February 1979. Later indications occurred on tubes in the row adjacent to the peripheral tubes.
4.3-4
TABLE 4.3-1 PLUGGED TUBES IN NUMBER 6 WEDGE AREA, B-STEAM GENERATOR, HOT LEG INSPECTION HISTORY AND CURRENT CONDITION EDDY CURRENT RESULTS AT PLUGGING DATE 1982 REVIEW CONDITION FOLLOWING JANUARY TUBE DATE 1982 EVENT. VIDEO TAPE (ROW-COLUMN) PLUGGED NEAR TUBE SHEET TOP SP 1 INSPECTION, VISUAL EXAMINATION 8-92 4/76 Very large"I.D. 2" above TTS Indication Collapsed and penetrated 9-91 4/76 Small I.D., not significant Dent Rippled O.D. surface 10-91 4/76 Small I.D. types ~ 3" above TTS Dent Wiped (Scraped) O.D.
11"91 4/76 Many I.D.s at TTS Dent Collapsed and tom open 4-5" above TTS 12-91 4/76 Large I.D. indications 4" Dent Collapsed and penetrated (LEAKER) above TTS hole at 4" above TTS 13-90 4/76 TS distortion Dent O.D. damaged 14"90 4/76 Large I.D. 4" above TTS, Small dent Collapsed and dented TS dent signal distorted 15-90 4/76 Large I.D., 3" above TTS, Distorted Collapsed rippled, and smaller I.D.'s penetrated 15-89 4/76 No confirmed indication Distorted 16-89 4/76 No indications Distorted Minor O.D. damage 17-89 4/76 Large I.D. 3" above TTS Clean O.D. damage Notes: TTS = Top of Tube Sheet SP 1 = Support Plate No. 1 I.D. = Inside Diameter T~e of Signal O.D. = Outside Diameter
TABLE 4.3-2 PLUGGED TUBES IN NUMBER 4 WEDGE AREA, B-STEAM GENERATOR, HOT LEG INSPECTION HISTORY AND CURRENT CONDITION EDDY CURRENT RESULTS AT PLUGGING DATE EDDY CURRENT RESULTS CONDITION FOLLOWING JANUARY, TUBE DATE AT INSPECTIONS PRIOR 1982 EVENT, VIDEO TAPE (ROM-COLUMN) PLUGGED NEAR TUBE SHEET TOP SP 1 TO PLUGGING INSPECTION + VISUAL EXAMINATION 45-46 2/82 41$ O.D. 24" above TS Clean None OK 45-47 2/82 49$ O.D. 24" above TS Clean None OK 45-48 4/80 1 O.D. - ~ 20" above TS Clean None OK 45-49 4/80 1 O.D. - ~ 20" above TS Distorted None OK 45-50 4/78 No significant 1/78 1 I.D. at TTS OK indication 45-51 4/78 1 I.D. at TTS 1/78 1 I.D. at TTS Mottled (rippled) O.D. surfa 45-52 4/78- Small I.D.s; possible Clean 7/77 Tube OK, *Removed in 4/78 bulge at TTS, one I.D. 1/78 1 I.D. at TTS about 5" below TTS 45-53 1/78 Many I.D.s, large Clean 7/77 Large, saturated Collapsed, O.D. damage amplitude, TTS = signals at TTS bulge signal 45-54 7/77 Large I.D. signals; Distorted 4/77 1 I.D. above TTS Collapsed, O.D. damage TTS location = ? 1 I.D. below TTS Severed at SP 1 possibly due to a bulge SP1 = Clean Missing 44-52 4/78 No indication Clean 44<<53 4/78 Unidentified signal Clean 1/78 Tube OK at TTS Sheet 1 of 3
TABLE 4.3-2 (Continued)
EDDY CURRENT RESULTS AT PLUGGING DATE EDDY CURRENT RESULTS CONDITION FOLLOWING JANUARY, TUBE DATE AT INSPECTIONS PRIOR 1982 EVENT, VIDEO TAPE (ROW-COLUMN) PLUGGED NEAR TUBE SHEET TOP SP 1 TO PLUGGING INSPECTION + VISUAL EXAMINATION 44-54 1/78 1 I.D. near TTS Clean 4/77 Tube OK Severed at TTS 7/77 Large, ~ 100$
indication at TTS 44-55 7/77 Large I.D. signals; Distorted 4/77 OK Collapsed, O.D. damage,-
TTS location = 7 severed at TTS .
possible bulge 44-56 7/77 Large, saturated Distorted 4/76 SP1 = Clean Collapsed, O.D. damage I.D. signals Dent between TTS severed at SP1 missing possible bulge and SP1 4/77 Tube OK, SP1 Clean 44-57 7/77 Large I.D. signals Clean 4/76 TS Distortion Collapsed, O.D. damage at TTS, possible bulge signal O
44-58 7/77 Large I.D. at TTS Clean 4/76 O.D. but TS Collapsed, O.D. damage possible bulge distortion 4/77 Tube OK 43-53 4/81 1 O.D. signal Distorted 43-.54 12/79 100$ signal above Clean TTS + Dent "at TTS 43-55 2/79 Possibly a small dent Distorted Severed at TTS or tube sheet distortion at TTS 43-56 2/79 O.D. signal 3" above Distorted TTS and large I.D.
signals at TTS 43-57 12/79 Large O.D. signal Dented 2/79 1 I.D. at TTS above TTS and TS SP1 = Dented distortion Sheet 2 of 3
TABLE 4.3-2 (Continued)
EDDY CURRENT RESULTS AT PIUGGING DATE EDDY CURRENT RESULTS CONDITION FOLLOWING JANUARY, TUBE DATE AT INSPECTIONS PRIOR 1982 EVENT, VIDEO TAPE (ROW-COLUMN) PLUGGED NEAR TUBE SHEET TOP SP 1 TO PLUGGING INSPECTION + VISUAL EXAMINATION 43-58 4/78 1 insignificant OK I.D. -1" above TTS 43-59 4/78 Very small I.D., Dented O.D. damage possible bulge, TTS 43-60 4/78 Small I.D., 5" above Dented 7/77 Tube OK, SP1 = Clean Scraped on O.D.
TTS, possible bulge at TTS 43-61 4/78 Nothing detectable Dented Severed at TTS 42-55 2/82 Burst 4.1 inches. x 0.75 inches 4/81 <40$ Axial O.D. Burst Indication; SP1 =
Distorted NOTES - TTS = Top of Tube Sheet SP 1 = Support Plate No. 1 I.D. = 'Inside Diameter +pe of Signal O.D. = Outside Diameter
- = R45 C52 was pulled in April 1978 Sheet 3 of 3.
C60 C59 C58 C57 C56 C55 C54 C53 C52 C51 C50 C49 ~ C48 C47 C46
/
R45 5'5*, '0 0 OO R44 3 3 55QQ Q O'E'42 Qa R43 0 0 ~
Burst 1/25/82 R41 INSPECTION 6 PLUGGING DATES 1 4/76 2 4/77
+ 3 7/77 4 I/78 A, 8 ~ Oate 5 4/78 Observed'.Y,Z + 6 2/79 OLINN 55 7 12/79
~ Bate Plugged 8 4/80 IO TYPE 00 OR NONE Type of Indication i~
10 9 4/81 2/82
~ Indicates possible Bulge accompanying ID - Type Signals 8 STEAN GENERATOR HOTLEG. NUHBER 4 HEDGE AREA: PLUGGING HISTORY Figure 4.3-1
4.4 Metallur ical Review An .extensive metallurgical evaluation is currently in,progress on a number 'of tube specimens which have been removed from the Number 4 and 6 wedge areas of the B-Steam Generator.
of these investigations is to provide documentation on whatpurpose The caused R42 C55 to burst, what are the failure mechanisms in the damage to tubes, and to help clarify why a number involved of periphery tubes have been plugged in the B-Steam Generator since April 1976. The results of the investigations to date are presented in Appendix A. The following is a synopsis of those results.
Extensive investigations have been conducted to date of 5 tube segments from the Number 4 wedge area. These include tube R42 C55 (which burst on January 25, 1982), R44 C54, R43 C54, R43 C55, R44 C55, and R43 C56. The techniques have included nondestructive examinations, dimensioning, macrophotography, optical metallography, scanning electron microscope (SEM), fractography, and microhardness determinations. In general, all degradation processes observed are exclusively mechanical in nature and are not in any way related to any properties of the Inconel 600 tubing material or to any environmental degradation of a corrosion nature. The physical properties of the tubing sections are normal for Inconel 600 and no evidence has been found of any stress corrosion cracking, pitting, or corrosion wastage. The elemental analysis of tube R44 C55 was typical of Inconel 600 (for Pe, Ni, Cr), and both this tube and tube R42 C55 gave mechanical property values which are typical of virgin, mil-annealed Inconel 600 (normal hardness, yield strength, and elongation).
Tube R42 C55, which suffered an axial burst, was shown to be worn axially in two long wear scars, one of which reduced the original 0.050 inch wall thickness to approximately 0.008 inches for approximately 4 inches in length. The resultant burst at this thinned point was a purely ductile failure. The fracture face was pure shear (dimpled rupture), the normal fracture mode for tensile overload in annealed Inconel 600. The burst is explicable exclusively in terms of a reduction (by wear) to a thickness which was insufficient to sustain the I.D.-to-O.D. differential pressure.
Severed, previously plugged tubes, R44 C54, R43 C55, and R44 C55, were shown to have been worn through in several areas. Fractures on these tubes were, in two cases, of a fatigue mode, the SEM fractographs revealing typical situations. This breakage of axial fractures is consistent with a wear process havingprocess reduced the"remaining axial ligaments to the point where these ligaments did fatigue and tear. Some of the other areas of breakage were not discernable, due to the fact that the fracture faces were obliterated by the damage process to the extent that details were lost to the SEM investigation. However, the evidence does support the fatigue or tensile overload of minimum ligaments which appear to-have operated to complete the damage process.
4.4-1
Tubes R43 C54 and R44 C56 'both showed wear areas consistent with those seen on the burst tube R42C 55. These tubes had been previously plugged. R43 C54 had leaked prior to plugging and had sustained further damage,:opening up a rather large area, due to wear and tube impacting.
All five tubes removed from the Number 4'wedge area showed wear areas. These areas of wear were axial, several inches in length, and exhibited circumferential grinding marks or striations. The wear areas were shown by metallography to have cold worked surface layers. These were verified by taking microhardness traverses across the cold work areas. Typical cold work on the wear surfaces was less than 0.001 inches in depth. Other wear areas, such as the burst wear surface, does not show any evidence of cold work.
Tube R44 C55, which was collapsed on the lower two inches of the specimen, was also analyzed for cold work by metallography and microhardness traverses. The results of these analyses documented extensive cold work (up to 0.008 inches in depth) in the collapsed area. A comparative analysis done at Battelle Columbus Laboratories on a specimen from tube R45 C54 also showed extensive cold work in the collapsed area of that tube..
Damage, for the most part, to both plugged tubes and to the burst tube R42 C55 appears to have originated in wear from the secondary side (O.D.). The worn surfaces then suffered additional degradation from tensile overload in the case of the active tube, and from fatigue and/or overload in the case of plugged tubes. Attempts to align the wear surfaces on the burst tube with wear surfaces on adjacent, previously plugged tubes demonstrated that the adjacent tubes were relatively free to move since they had been severed near the top of the tube sheet. On the burst tube, for example, the wear scar associated with the burst faced the backside of the neighboring tube R43 C55, and faced the wrapper. There was, however, a second wear scar on R42 C55 which was displaced circumferentially by an amount which confirms that severed tube R43 C55, which caused the wear, was mobile. Examination of other worn, previously plugged tubes, also revealed multiple wear scars (axially oriented), suggestive again of the mobility of the severed wearing members.
In addition to these results for tubes from the Number 4 wedge area, which are detailed in the appended report, examinations are continuing on three tubes from the Number 6 wedge area. Zones which contained the 1976 eddy current indications have been radiographed on tubes R12 C91 (the 1976 leaker), R14 C90 and R15 C90.- .No evidence of I.D. anomalies or discontinuities was discernible from the radiographs. Additionally, the destructive (metallographic) examination of the I.D. surface of tube R15 C90 in the zone from 1 to 6,inches above the tube sheet has not revealed any characteristics which could account for the 1976 type of eddy current signals. These laboratory in progress to verify the structural damage these tubes investigations'emain suffered as well as their original plugging defect origins.
4.4-2
4.5 Anal ysi s This section describes the results of analysis progr ams performed to establish the nature and magnitude of loads acting on steam generator tubes; and evaluates analytically the role that these loads played relative to the postulated failure mechanism described in Section 4.2.
Consistent with the observed degradation of tubing in the 8-Steam .
Generator, the tubing of primary interest is the tube spans between the tube sheet (TS) and the first tube support plate (TSP) in the neighborhood of the wrapper support blocks on the hot leg side of the steam generator. Accordingly, where applicable, the localized conditions associated with this tubing were utilized in the evaluation.
Figure 4.5.1-1 shows schematically this region of the bundle. The figure also includes the pertinent dimensions of the tube and interacting components.
The analyses included thermal mechanical loads due to normal operating transients, lateral loads due to a foreign object impact, and hydraulic loads including consideration of vortex shedding, cross-flow turbulence and fluid-elastic excitation. Axial and bending loads were included in the evaluations to account for axial tube restraint at the TSP and possible offset between the TS and TSP holes. If tubes are axially restrained at their ends (tube sheet and TSP), axial loads can develop due to thermal growth mismatch of a plugged tube relative to the stub barrel and/or surrounding active tubes. Axial bending stresses result from (1) an initial offset between the TS and TSP holes and radial thermal growth mismatch between the TS and TSP, and (2) rotation or bowing of the tube sheet due to the primary-secondary p essure differential. The initial maximum offset was assumed to be 0.1 inch based on the fabrication practice and experience with Model 44 and similar design steam generators. These'axial loads were developed conservatively assuming full restraint of tubes in the wedge areas; away from the wedge areas, axial restraint would be significantly lower due to the TSP flexibility.
Following the load determination, a structural evaluation of tubing.
subjected to these loads is presented considering the potential of collapse, fatigue, wear and burst.
4.5.1 Oesi n Parameter s I
Steam generator and tubing design parameters, and transient deScriptions are presented in the following:
4.5.1.1 Steam Generator No load. steady-state. conditions Thot Tcold = Tav = Tstm = 547 F Pp 2250 psia Ps = 1020 psia
- 4. 5-1
100. P ercent. Ful l-Power Condition Thot = 603,8 F Tcold = 547 7 F Tay = 572.5 F Tstm = 516 ~ 3 F Pp = 2250 psia (hot leg)
Ps 777 psia TFW = 414.7 F Steam flow = 3.1 x 106 lbs /hr V
Feedwater flow = 3.18 x 106 lbs /hr Recirculation ratio = 4.73 Tube Bundle Par ameters Pitch = 1.234 inches Wrapper I.D. = 117.00 inches Wrapper opening = 14.00 inches Tube span between tube sheet and first support plate = 51.81 inches Ferrous Materials and Pertinent Properties Tube sheet SA-336 Stub Barrel SA-302 GR B Support plate, shear lugs, spacers and wedges Car bon Steel E70* = 27.9(10)3 ksi E600* = 25.7(10)3 ksi a70*+ 6 7(10 in- F a<o ~~= 7.23(10) in- F oungs o u us at the .indicated temperature
- = mean coefficient of linear expansion in going from 70 F to the indi-cated temperature.
4.5-2
Steam generator tubing material is mill-annealed Inconel-600. The nominal tube dimensions are 0.875 inch O.D. x 0.050 inch wall. The design minimum wall is 0.045 inch and the maximum specified ovality in the straight legs of the tubing is two percent.
Pertinent material properties for the evaluation are listed below:
E70 = 31.7(10)3 ksi E600 = 29.2(10)3 ksi a7 -- 7.18(10) a600 7 90 10 in- F Sy+ = 35.5 ksi at 550 F**
Sa = 26.0 ksi = Section III fully reversed alternating stress limit at 106 cycles for E = 26(10)3 ksi 4.5.1.3 Transient Descri ption In the following transient description, active tubes are. assumed to
.interact with both primary and secondary fluid; plugged tubes are considered to interact only with the secondary fluid and as such are assumed to be in thermal equilibrium with the secondary fluid.
a) Plant Heat-Up (200 cycles)***
This transient is initiated from the cold shut-down condition. At the end of the transient the primary pressure and temperature are 2250 psia and 547 F, respectively. The heat-up rate is such that continuous thermal equilibrium is maintained between the primary and secondary fluids. Accordingly, at the end of the transient, the secondary fluid is at the no-load (hot standby) steady-state temperature of 547 F and the secondary pressure equals the saturation pressure, 1020 psia.
atersa yse strength
- This is the code minimum value. Acceptable range at room tem-perature for yield strength is 40.00 65.00 ksi per applicable specifications.
- Numbers in parenthesis are design basis cycles for 40-year plant 1 i fe.
4.5-3
b) Plant Cool-down (200 cycles)
This transient is an exact reverse of the plant heat-up transient in a) above.
c) Plant Loading (14,500 cycles)
This transient initiates from hot standby conditions. The primar y pressure during this transient remains at 2250 psia, while Thot increases from 547 F at no-load to 604 at full load. The secon-dary side temperature and pressure drop to 517 F and 780 psia, respectively.
d) Plant Unloading (14,500 cycles)
This transient initiates from 100 percent full power conditions and is an exact reverse of plant loading transient in c) above.
e) Step load transients of + 10 percent full power (4,000 cycles).
For fatigue evaluation, these transients are conservatively enveloped by the plant loading/unloading transients.
Reactor Trip (400 cycles)
For fatigue evaluation, this transient is enveloped by the plant unloading transient.
g) 50 percent Step Load Decrease (200 cycles)
Also enveloped by the plant unloading transient.
h) Hot Standby with 70 F Feedwater Cycling (200 cycles)
This transient deviates from the Design Specification in that for Ginna plant, since early 1980 the feed is continuous at a rate of about 35 gpm (as opposed to a cyclic slug-feed at 300 gpm in the Design Specification). Although normally only one occurrence per year is expected for this transient (during plant start-up following refueling), for analysis purposes, 200 cycles of the transient are conservatively assumed, consistent with the number of heat-up cool-down transients.
4.5.2 Thermal-Hydraulic-fvaluation 4.5.2.1 ~Per se The purpose of this sub-section is to determine fluid velocities for use in evaluating fluid induced loads on steam generator tubes near the tube bundle entrance region.
The plugging of tubes in the periphery of the tube bundle will lead to a redistribution of the flow in the plugged tube region. The flow velocities resulting from this redistribution must be evaluated to 4.5-4
determine if flow induced loads on the steam generator tubes in and near the plugged region are significantly affected.
The CHARM computer program (reference 4.5.2-1) was used to perform a thermal-hydraulic analysis of the tube bundle region between the tube sheet and the first support plate. Two cases werd considered: a) nominal, and b) one block of tubes plugged in the periphery of the bundle.
The two-dimensional CHARM analysis was performed for the plane of symmetry perpendicular to the tube lane which divides the hot and cold legs into equal halves. At nominal conditions, there is no flow across this plane.
Because the CHARM analysis was two-dimensional (axial and radial) a separate three-dimensional hydraulic analysis of the tube sheet-to-first support plate region was performed to determine circumferential changes in the fluid flow patterns due to plugging. This analysis used the WECAN hydraulic conductance element (reference 4 .5 .2-2) and a pressure-forced boundary condition, rather than the flow-forced assumption inherent in the CHARM analysis.
4.5.2.2 CHARM Analysis CHARM, is a Westinghouse Proprietary two-dimensional analysis code used to compute the fluid flow conditions in a two-dimensional domain. The basic variables computed are the pressure, velocity, density and enthalpy by use of multidimensional characteristics in whi ch a system of fluid flow compatibility equations are formulated and numerically integrated along bi characteristic curves. The region analyzed is divided into rectangular cells of sizes sX and aY along the two (X, Y) coordinate axes. Velocity and pressure at the boundaries are specified and the code calculates the pressure, velocities (lateral and vertical),
temperature, density, enthalpies, and quality distributions inside the model. Other input data required are the flow resistances (either input or calculated internally by the code) and heat load in each cell of the model.
Since the code does not recognize the presence of obstructions inside a
- cell, the tubes in the model were simulated by flow resistances. The input velocity to the model was adjusted to account for the reduction of the free volume available for the fluid as a result of tubes being present. The output velocity was then corrected to restore its physical meaning.
4.5.2.3 RGE.Steam Generator. Model.and In ut Data The region modeled for the CHARM calculation was the area between the tube sheet and the first tube support plate (1st TSP). The plane of analysis was the vertical diametral plane perpendicular to the tube lane axis of the steam generator. The model, therefore, consisted of the hot side and the cold side of the unit wi th the centerline of the SG located at the middle of the model plane as shown in Figure 4.5.2-1.
4.5-5
The tubing inside the cells of the model was represented by radial and axial flow resistances. The input data were the tube O.D. and pitch.
The numerical values of the resistances were computed internally by the code. When the tubes in a certain area were plugged, the heat load in the cell was set to zero.
Relevant geometrical data used in the CHARM calculations were:
Tube O.D. 0.875 inch Tube Pitch = 1.234 inches Wrapper I.D. = 117 inches.
Wrapper Opening = 14 inches Tube sheet-to-First Support Plate Height = 51.81 inches The flow inlet velocity at the wrapper opening on the hot side was taken to be 2.38 ft/sec and that on the cold side to be 2.14 ft/sec. These velocities were consistent with the split of the total flow between the hot side and the cold side at the tube bundle inlet being taken to be
.65 and .35, respectively, which accounted for some mixing and swirling in the downcomer annulus.
As mentioned previously, the code does not recognize the porosity of a cell. As a result, the input velocity to the boundary cells of the model was adjusted to account for a reduction in area due to the presence of tubing using the appropriate tube pitch and diameter.
4.5.2.4 Three-dimensional WECAN) Analysis The WECAN hydraulic conductance analysis model was based on the subdivision of the tube sheet-to-first support plate region into the cells shown in Figure 4.5.2-2. The extent of the plugged tube region was assumed to be five tube pitches deep (radially) a'nd 20 wide (circumferentially). This corresponds to approximately 80 plugged tubes or more than twice the actual number of plugged tubes in the Number 4 wedge area. The model consisted of 100 internal nodes, each at the center of one of the cells shown in Figure 4.5.2-2. Five boundary nodes (201, 206, 211, 216, and 221) were used to represent the fluid just upstream of the wrapper opening. Twenty-five boundary nodes (101-125) were used to represent the boundary at the first support plate.
Figures 4.5.2-3 and 4.5.2-4 show the nodal distribution in the e = 10 and Z = 7-inch planes, respectively. Node numbers in the remaining e and Z planes were such that the node-spacing in the circumferential direc- tion was 5 and that in the axial direction was 25. In the radial direction, the node spacing was made smallest in the vicinity of the region where the tubes were plugged, and varied in a geometric progres-sion radially inward. The circumferential spacing of nodes was a constant 20 and allowed for two nodes per axial level on each side of the sector in which tubes were plugged.
The WECAN hydraulic conductance model calculates the flow rate between any two nodes according to the formula:
- 4. 5-6
e FLOW RATE = C X J a p where C is the hydraulic conductance and hp is the difference in pressure between two nodes. The hydraulic conductances between internal nodes were determined from CHARM calculated cross-flow and axial flow friction factors according to the following:
- ~
H g~
C (axial) = A ax 2DTg C (circumferential = A and radial) cr where p. fluid density (.7812 x 10-4 lb-sec2/in 4)
DH = axial flow hydraulic diameter (1.3406 in)
DT = tube D.D. (.875 in) fax = axial flow friction factor = .03095 fcr = cross flow friction factor = .99043 g = 386 in/sec2 L = distance between nodes A = flow path area between nodes In the wrapper opening flow paths (nodes 201-1 through 221-21) the hydraulic conductance included both the tube crossflow component and the hydraulic loss due to the expansion and turning of the flow as the downcomer through the wrapper opening. To simulate the low fluid it exits velocities indicated by CHARM in the plugged-tube region above the wrapper opening, the axial hydraulic conductances in this region were set equal to very small numbers.
Calculations were performed by setting the pressures in the downcomer nodes (201-221) equal to .96 psi and those at the first support plate (101-125) equal to 0.0 psi to approximately simulate the appropriate pressure drop in this region. By pressure-forcing, rather than flow forcing the calculations, the circumferential flow velocities can be estimated. The downcomer and first support plate were chosen as boundary points for pressure-forcing because the pressure at each of these two points should be relatively constant.
4.5.2.5 Resul ts The CHARM calculations included the following cases: a) Ginna nominal base case at 100 percent power, and b) a perturbed case simulating a plugged tube region in which the heat load in all cells along the columns X=2 and X=20 were set to zero. The widths of these columns are 5.85 inches, or approximately five tube pitches.
Figures 4.5.2-5 and 4.5.2-6 show the CHARM computed average velocities in the cells for the nominal and plugged cases at the hot side
.entrance. The interpretation of the output is as follows:
4.5-7
- 1. To obtain the physical lateral velocity (UL):
Tf 2 p2 gD L CHARM p
where UCHARM is the CHARM computed velocity.
- 2. To obtain the physical axial (or vertical) velocity V: No change
- 3. To obtain the maximum lateral velocity between tubes (U ax gap):
lf 2 p' --V.D- ~ .p ~
max, gap CHARM p2 P D Figures 4.5.2-7 through 4.5.2-10 show the velocity vector and quality distributions for both cases. The major effect of tube plugging was the appearance of a reduced fluid velocity field and a low quality region in the area between the bottom of the wrapper and the first support plate where the tubes were plugged. Because the density of the fluid in this region was lower, the resulting pressure gradients tended to be such that the fluid entering the tube bundle through the wrapper opening does not expand as quickly as it did in the nominal case. This tended to incr ease fluid cross flow velocities in the plugged-tube region near the wrapper opening.
In the CHARM model, circumferential fluid velocities were neglected.
Figure 4.5.2-11 illustrates the behavior of the flow field in the plugged-tube region near the wrapper opening (node by the 3-D WECAN analysis.
ll cell) as predicted Because of the flow restriction produced by the deadwater region above the wrapper opening, the fluid enter ing the plugged-tube region expanded circumferentially (face C) instead of axially. The flow restriction also resulted in a reduction of the incoming velocity on face A. On face B, the velocity was below the value because of the efflux through faces C. 'HARM-predicted Conservative estimates of the limiting cross-flow velocities required that the significant effects of both the CHARM and WECAN analyses be considered. The major result from the CHARM analysis was an increase above nominal values of the radial velocities in and near the plugged tube region. The major results of the WECAN analysis were the predic-tion of a reduced flow out of the plugged tube region (face B) and of a circumferential crossflow velocity component (between tubes) of approxi-mately 3.5 ft/sec (through face C). A limiting location on faces A and C, therefore, was their intersection, where the effects of both radial and circumferential components were p.esent. On face 8, however, more conservative to take the CHARM-predicted value rather than the it was WECAN value, because the latter is significantly lower due to the efflux faces C. The CHARM-value on face B, however, will be, in 'hrough addition, conservatively adjusted to reflect the fact that the area of face B is less than that of face A. Mathematically, these maximum cross flow velocity estimates are summarized below:
- 4. 5-8
2 +
Face A: UCR
= U UC2 n0m r"' 'uc =UU CHARM area of face A Face C: Same as on face A UCR where UCR is the cross flow velocity and UpAR" is the CHARM velocity at face B.
Table 4 .5 .2-1 sumlarizes the results of these calculations for nominal and plugged tube conditions. Circumferential crossflow velocities were used only in the plugged tube cell; calculations indicated that circum-ferential velocities elsewhere are negligible. The face B crossflow velocities are all CHARM-predicted values adjusted for the area reduc-tion. Also given are adjusted CHARM crossflow velocities on face B for several cells downstream (radially inward) of the fir st cell. The nominal between-tube crossflow velocity is 8.2 ft/sec and occurs at the outer periphery of the tube bundle.
In summary, the p esence of a block of plugged tubes approximately five tube pitches deep (approximately 80 tubes) had the following effects:
a) A reduced fluid velocity field and a low quality region appeared between the wrapper opening and the first support plate in the region where the tubes are plugged. This tended to increase fluid crossflow velocities in the plugged tube region near the wrapper entrance.
b) The highest between-tube crossflow velocity increased from 9.01 ft/sec in the no-plugged tube case, to 9.11 ft/sec in the plugged tube case, or slightly more than one percent, which is not significant.
4.5e3 Axial.Roads The development of axial loads as set forth herein is limited to plugged tubes since evaluations which follow that utilize this information focus on plugged tubes. Axial loads on steam generator tubing are dependent on the degree of tube-to-TSP restraint and could result from any of the following three considerations:
<) External Pressure: If the tube is free to move axially through the TSPs, a net compressive load is developed on the tube due to the difference in projected areas of the outer and inner U-bend surfaces. This load is a maximum at no-load condition. The magnitude of this load is F = m/4 Do2 Ps 604 lb (compression) 4.5-9
where Do = 0.875 inch, outside tube diameter, and Ps 1005 psig, maximum secondary side pressure b) Tube-to-Stub Barrel Thermal Interaction: If a plugged tube in the wedge area is axially restrained at the TSP, axial loads in the tube result due to the differences in stub barrel/shell and tubing thermal displacements. The range of axial load on a plugged tube in a wedge area on hot leg is calculated for the significant transients and full power steady-state conditions.
Hot. Standby with 70 F Feedwater Minimum shell temperature for this transient was calculated to be 476 F assuming slug-feeding at the 300 gpm maximum rate per Reference Speci-fications*. The plugged tube was conservatively assumed to remain at the no-load steady-state temperature of 547 F. The maximum tube load is:
F = EaAt where E = 29.4 x 106 lbfin2, elastic modulus of tube lr/4 (Do Di~) = 0.12 in2, cross sectional area of tube, and e is the tube restraint per unit length due to thermal growth mismatch with the stub-barrel and is given by (a aT)SB = 7.67 x 10-6 x (467 547) = 0.00055 in/in maximum Thus, the maximum resulting tube load is F = - 1940 lb (compression)
This load reduces to zero as the unit reaches the no-load steady-state condition. During this transient, since only the minimal flow velo-cities exist, the load is not relevant for flow-induced vibration analyses, but should only be considered in fatigue evaluation.
P l ant. Loadi n /Un] oadin For these transients, the mean shell temperature was conservatively assumed to remain at its corresponding value at the initiation of the transient. The plugged tube temperature, however, closely follows the actual secondary fluid temperature. Consequently, at the end of the plant loading transient, the tubing is subjected to an axial tensile load due to the temperature differential of aT = 547 517 = 30'F. Or, F = E (e hT)SB At = 825 lb (tension)
- The Ginna procedure since early 1980 has utilized continuous feed at a nominal rate of 35 gpm, and results in significantly less severe transient conditions than analyzed here.
4.5-10
Upon reaching steady-state full power conditions, the mean shell tem-perature is about 20 F lower than the bulk secondary temperature due to subcooling between the wrapper and the shell. Therefore, even though the plant unloading transient is an exact reverse of the plant loading transient, the tube is subjected to an axial comp"essive load corresponding to zT = (30 + 20) F, or F = 50/30 x 825 = 13?5 lb (compression)
Full Power. Steady-State As explained above, the tube is under a constant compressive load corresponding to hT = 20 F due to the subcooling effect on the mean shell temperature. Hence, F = 20/30 x 825 = 550 lb (compression) c) Plu ed Tube-to-Active Tube Interaction: If a plugged tube is surrounded by a cluster of active tu es, and if the tubes are axially restrained at the TSP, the plugged tube motion is essen-tially controlled by the active tubes. The load in a plugged tube that exists at 100 percent power condition can be calculated as follows:
Assuming no thermal loading at the no-load, uniform steady-state conditions of T = 547 F, the difference in thermal growth is:
< = (ahT)A (ahT)P = 0.0004 in/in where mean metal temperature variations for the active (A) and plugged (P) tubes and linear coefficient of thermal expansioh are:
hTA = 565 547 = 18 F hTp = 517 547 = -30 F and a = 7.80 x 10-6 in/ in The resulting force (tension) on the plugged tube is:
F = EtlAt = 1410 lb Summary In addition to the above conditions, axial loads due to other transients can be calculated in a similar fashion once the temperature differen-tials between the tube(s) and stub barrel/shell are established. In using these loads for the various analyses, the following should be noted:
o Axial loads due to pressure and thermal growth mismatch cannot occur simultaneously on a given tube, as these loads result from indepen-dently assumed tube end conditions at the TSP.
4.5-11
4 o At the no-load, steady-state condition, all components are at a uni,form temperature of 547 F and, hence, axial tube loads following the hot standby and plant unloading transients approach zero as the temperatures approach steady-state conditions. These loads are, therefore, applicable to fatigue analyses, but would have a minimal influence on collapse as a result of lateral impact loading, and flow-induced vibration analyses due to the short period of time they would be acting.
o, During full power steady-state conditions, the compressive axial load on a tube exists due to the subcooling effect on the mean shell temper ature. However, the load due to tube-to-tube interaction is significantly higher, although tensile. The steady-state compressive load is important from the viewpoint of increased susceptibility to fluid-elastic instability; and the tensile load is important from the viewpoint of increased susceptibility to collapse.
Maximum calculated axial loads for a wedge area on hot leg and their analytical applicability are suoearized in Table 4.5.3-1. Axial loads in a nonwedge area would be less.
4.5.4 Lateral Loads 4.5.4.1 ~Pur se Lateral loads on steam generator tubes in addition to a pressure dif-ferential between the 0.0. and I.D. surfaces can result from interaction of the tubing with an external source such as the fluid or a solid object. Loading due to fluid effects will be discussed in 4.5.6 and the affects of a pressure differential is well established. The purpose of this sub-section will be to estimate analytically the magnitude of the lateral impact load that a solid object, similar to that removed from the Ginna steam generator could input to a tube.
4.5.4.2 Analysis A schematic of the tube bundle entr ance region with a foreign object present is shown in Fi gur e 4.5.4-1. Fluid forces acting on the solid object can accelerate it toward a tube resulting in impact and tube motion. After coming to a halt, the elastic forces of the tube then force the foreign object in the direction opposite to that from which it came and disengagement of the solid object and tube can occur. If this happens, one or more additional impacts can occur. The analysis of this phenomenon proceeds as follows:
4.5.4.3 FFi3.35 Mhen the foreign object is disengaged from the tube the force acting on it from is strictly hydrodynamic. The magnitude of this force is estimated F (f'luid) = CU JFr AIUr XPI(uf xP)/2 wher e 4.5-12
Pf ~ = fluid density = .7812 (10-4) lb-sec2/in4 A = projected area of loose part (assumed to be a triangle 6.5 in by 4.19 in).
Uf = Fluid Velocity xp fore i gn ob ject vel oci ty CD = Drag Coefficient The dr ag coefficient is estimated from Reference 4.5.4-1 to be equal to unity. The equation of motion of the foreign object is then Mp x p F ( f 1 ui d) where M>
foreign object mass (assumed to be .0074798 lb-sec2/in),
and 'xp ss the foreign object acceleration.
The fluid velocity Uf is assumed to have the following form:
Uf = Uss + UA cos (2 m ft) where Uss = A steady state. component of velocity UA = Al ternating component of velocity f = Frequency of alternating velocity component t = time The reason for including the alternating velocity component is to simulate the type of alternating velocity fluctuations that would be expected downstream of a support block in the downcomer. It is estimated that, for a 6.5 inch wide support block in a 15.5 ft/sec stream and a Strouhal number of 0.2, the frequency of such fluctuations is approximately 5 .72 HZ.
4.5.4.4 Steam Generator Tube The steam generator tube is simulated as a simple spring-mass system initially at rest. Two spring-mass combinations are considered. In the first model (Model 1) the deformation of the tube is assumed to follow the mode shape of the first fixed-fixed beam mode between the tube sheet and the first support plate. As such the applied loading is related to the mass per unit length of the tubing between the tube sheet and first support plate. The spring constant based on a deflection four inches above the tube sheet is 3717 lb/in and the effective mass is .03767 lb-sec2/in. In the second model (Model 2), the spring constant and mass are 19439 lb/in and .01068 lb sec2/in, respectively, and are based on the static deflection profile of a fixed-fixed beam with an applied point load four inches above the tube sheet. This second model is more representative of a lateral impact load being applied to a tube such as occurred at Ginna. Shell stiffnesses have been neglected.
4.5-13
The tube equation of motion is obtained from Newton's Second Law using the nomenclature of Figure 4.5.4-1 and is:
(MSG + E Mp) x + KSG x = 0 I
where E = 0 corresponds to the situation when tube and foreign object have independent motions, and E = 1, to the situation when they move to gether.
Both foreign object and,tube velocities at time zero are set equal to zero, as is the initial position of the tube. The foreign object and tube equations of motion are numerically integrated in time.
4.5.4.5 Results Figures 4.5 .4-2 through 4 .5 .4-7 illustrate the Model 1 steam generator tube force history, and the displacement histories of both the steam generator tube and foreign object for three cases. These cases are summarized below:
Case Xp (0) Uss UA
-5 inches 2.3 ft/sec 0.0 ft/sec 0.0 inches .69 ft/sec .69 ft/ sec 0.0 inches .345 ft/sec .69 ft/sec Case 1 considers the translation of the foreign object from a position near the shell to the steam generator tube as a result of fluid forces associated with a nominal fluid velocity of 2.3 ft/sec. Cases 2 and 3 simulate the type of behavior which might be expected, when the steady fluid velocities are lower than nominal and alternating velocities exist because of a disturbance produced by an object like a support block.
The alternating velocity was assumed to be 30 percent of the nominal steady-state value, or .69 ft/sec Foreign object induced loads associated with each of the above cases are provided in Table 4.5.4-1. The impact forces are on the order of 26-107 lb in Case 1. In cases 2 and 3, it is notable that the impact forces are higher for the lower steady-state velocity (Case 3). This is because disengagement and impact can more readily take place when the steady hydrodynamic forces are small.
A comparison of the Model 1 and Model 2 results demonstrates the importance of the tube dynamic characteristics on the magnitude of the impact loads. In Case 1, the magnitude of the initial impact load will be approximately proportional to the square root of the ratio of the spring constant to the total mass (tube-effective plus foreign object) of the system. Consequently as verified by the analysis the Model 2 loads should be greater than the Model 1 loads because of the higher stiffness and lower mass associated with Model 2. The stiffness associated with impacting of a steam generator tube by a foreign object will be greatest near the top of the tube sheet because of the restraint provided by the tube sheet to a beam type deformation. If the primary 4.5-14
S tube response to impact is a shell type rather than a beam-type of deformation, the high stiffnesses and low effective-mass associated with a shell mode can be expected to yield impact forces even greater than those given in Table 4.5.4-1.
In summary, this analysis demonstrates that foreign object impact loads in excess of 100 lb are possible even for beam-type only deformations of the impacted tube. In all probability, shell-type deformations will also be significant, and will lead to even higher impact loads.
4.5.5 Col < apse The purpose of this section is to analytically assess the potential for tube collapse due to possible thermal, mechanical and hydraulic loads identified for the Ginna steam generator in conjunction with the postulated failure mechanism described in Section 4.2.
In this analysis, the tube was assumed plugged and thus was subject to the secondary side external pressure. In addition, an axial load which arose from restraint at the first tube support plate was assumed to act on the tube. A radial load associated with a foreign object impacting the tube near its tube sheet end and causing local deformation of the tube wall was then superimposed on the pressure and axial restraint loadings. Random and repeated application of such a load can result'in degradation of a local area of the tube surface and in progressive tube ovality such that, under the combined effects of the axial restraint, pressure, and impact, the tube would collapse when the ovality reaches a critical value.
In the following, a lower bound calculation using the von Mises cri-terion with only the external pressure and the axial load acting on the tube was performed first. Then, a finite element analysis of an ovalized tube is performed to include the effects of a radial concen-trated loading.
4.5.5.1 The von Mises Criterion One method that is often used to determine yield in a structure is the
. von Mises criterion. It is essentially a measure of the maximum dis-tortion energy. For a cylindrical tube under external pressure and axial loading, the von Mises criterion may be written as:
2 Sy (zh-oa)2 + (+a ar)2 + (zr ah)2 Her e:
Sy yield strength of the tube material ar -P, the external pressure Rm
+p~
oa F/2mRmt 4.5-15
where:
~a axial stress
~r radial stress hoop stress R.h mean radius of tube t tube wall thickness F total axial force Equation (1) states that one half of the sum of the squares of the stress intensities can be taken equal to the square of the material yield strength at incipient yielding of the structure.
If D = O.D. of tube, then R =
~ and Equation 1 becomes:
-P( -l)a a +P 2 t + )-S y =0 2 Rm Rm Rm-a a
(1+ (2) t2 For the Ginna steam generator tubing:
D = 0.875 in Rm/t = 8.25 t = 0.050 in A = m/4 (0.8752 0 7752)
Rm = 0.4125 in = 0.12959 in2 Thus, Equation 2 simplifies to 2 + 77.3125 2 2 =
cr a
(7.25) a a P S 0 (3)
Since aa is due to the axial force (F), Equation 3 can be used to determine values of F.
Equation 3 is plotted in Figure 4.5 .5-1 for Inconel tubing with yield strengths of 38 ksi and 55 ksi. For reference, the Euler's buckling load and a collapse test point due to external pressure only are indicated. It is seen that under usual values of pressure and axial force, the plugged tube with a nominal thickness will not collapse.
4.5.5.2 Stress Analysis-of-Ovalized Tubes The stress analysis of ovalized tubes subjected to combined external pressure, localized radial loading, and axial loading is described in this section. The purpose of the analysis is to determine the mag-nitude of radial load required to produce elastic stresses equal to the yield point of the material, when such a load is superimposed upon a constant external pressure loading of 1000 psi and, in some cases, a constant axial tensile load of 1000 lb*. The analysis, though related
- As discussed in Section 4 .5 .3 and Table 4.5 .3-1, during steady-state power operation, tensile loads of this order can be expected in a plugged tube near the hot leg wedge location due to the thermal growth mismatch under the assumption of axial restraint at the first TSP.
4.5-16
to the investigation of the collapse of plugged steam generator tubes, was not a true collapse load analysis. It only indicated the load at which yielding begins at the highest-stressed point. Prior to collapse, yielding must penetrate essentially through the thickness of the tube wall and far enough along the tube length to precipitate instability.
To account for this series of events would require a true collapse 'load analysis. Such an analysis, under these conditions would be much more extensive than the one described herein, involving complex modeling and iterative, large deflection, elastic-plastic stress calculations.
The analysis was performed for 0.875 inch 0.0. tubes of two wall thicknesses, 0.05 and 0.03 inch, and four ovalities, 0.01, 0.02, 0.04 and 0.06. The tube material is Inconel 600 at 550 F, with a minimum yield strength of 35 .5 ksi, an estimated maximum yield st< ength of 57.7 ksi and a modulus of elasticity of 29.4 x 10o psi.
Computer analysis was performed with the ANSYS program. The initial intent was to use a single structural model to handle both external pressure and localized loading. However, after expel iencing problems with boundary condition simulation for pressure loading, it was decided to use a separate supplemental model that had previously been success-fully employed in tube collapse pressure analysis. This model is shown on Figure 4.5.5-2. The model used for concentrated radial loading is shown on Figures 4.5.5-3 and 4.5.5-4. The effect of axial loading on stresses was superimposed upon computer results by means of hand calculations. The von Mises yield criterion was assumed to apply.
Results of the analysis are plotted on Figures 4.5.5-5 and 4.5.5-6.
Figure 4.5.5-5 plots maximum von Mises stress vs. concentrated radial load for the two tube thicknesses. In Figure 4.5.5-5, it is seen that the maxinam stress increases with increasing concentr ated radial load, increases with. increasing ovality, increases with decreasing tube wall thickness, and increases with application of a 1000 lb tensile load.
For the nominal tube wall thickness of 0.050 inch, it is seen that the maxitmm stress in the tube exceeds the maximum yield stress of 57.7 ksi at an ovality of 0.01 and a concentrated radial load of approximately 75 pounds. Geater tube ovaIity reduces the radial load required to speed the maximum yield stress of 57.7 ksi; for example, at an ovality of 0.06, the maximum yield stress is reached at a concentrated radial load of approximately 60 pounds.
Figure 4.5.5-6 is a plot of hoop stress versus axial distance from the point of load application for a five pound concentrated radial load.
This figure shows the localized nature of these stresses. Such localization would indicate that substantial yielding in the immediate vicinity of the load could probably be tolerated before'ube collapse would be precipitated. The first phase of yielding is visualized as a local, surface effect, with a subsurface elastic core and fully elastic tube sections a short distance away from the load continuing to provide stability. Only when the yield zone had spread through-wall and over a substantial area, will collapse occur. Typically, plastic hinges along two diametrically opposite planes for a length of 2 to 3 tube diameters are required for collapse. Thus, for eventual collapse in addition to a larger than analytically calculated magnitude of radial load to cause 4.5-l7
yielding, the radial load must act in an impacting'manner over a distributed area spanning about 2 to 3 tube diameters along the tube axis.
4.5.5.3 Factors-in the Co)la se of the Ginna Tube It is seen from the preceding sections that the Ginna tube can deform plastically under a combination of external hydraulic (secondary side) pressure, axial load and a concentrated radial load. The latter load can cause the tube wall to deform progressively with repeated appli-cations resulting in increased ovality and eventually, the external pressure will collapse the tube.
4.5.6 Flow-Induced Vibration teristicss The purpose of this sub-section is to determine the stability charac-and magnitudes of the flow-induced vibrational displacements and/or loads for hot leg tubes between the tube sheet and the first support plate. This information is pertinent to assess the degree of fatigue usage as a result of fluid-solid interactions. Both gross and local hydraulic effects will be considered. Gross hydraulic effects are defined as those which act on a tube or group of tubes and are of interest relat'ive to the overall response characteristics of a tube.
The loading mechanisms are fluid-elastic excitation, vortex-shedding and .
turbulence. Effects of tube plugging, structural degradation and axial loading are considered in the analysis.
In the local hydraulic load analysis, the magnitude and frequency of lift, drag, and torque loads acting on a tubing protrusion are cal-culated and are of interest in assessing the propensity for shredding of a tube due to fluid loading. The primary mechanism for this is con-sidered to be vortex-shedding.
4.5.6.1 Gross Fluid Effects In this sub-sub-section, results of analyses perfor med to characterize the dynamic behavior of tubes with cylindrical and irregular cross sectional geometries, when influenced by secondary fluid cross-flow, are reported. The purpose of the analyses was to determine tube natural frequencies, stability to fluid elastic excitation, and amplitude of vibration in terms of cross sectional geometry and cross-flow velo-cities. The analyses were based on an undistressed'cylindrical tube, and other tube configurations having flat, 10 percent ovalized, and kidney shaped distressed cross sections. Fluid elastic excitation, vortex shedding, and cross-flow turbulence were the vibration mechanisms evaluated for tubes with a fixed-fixed and fixed-pinned boundary.
4.5.6.1.1 Mathematical. Model of the Tube Confi urations Specific parameters used in the analyses are presented in Table 4.5.6-1, and the basic model configuration between the tube sheet and the first tube support plate is shown in Figure 4.5.6-1. Individual mathematical models were generated to simulate either an undistressed cylindrical tube or cylindrical tubes having a distressed cross section four inches above the tube sheet and two inches in length. The distressed cross 4.5-18
sections were: totally flat, 10 percent ovalized, and a kidney shape.
Since the tube was considered to be plugged, the total mass of the tube was the summation of the tube material mass and the mass of the secon-dary fluid displaced by the tube. Fixed-fixed boundary conditions constrained all degrees-of-freedom at the tube sheet and the first tube support plate. The fixed-pinned boundary constituted total fixity at the tube sheet, and constrained tube displacements at the first tube support plate, where rotation in the x-y plane was permitted. Natural frequencies and fluid-elastic stability ratios and amplitudes were computed using a lumped mass finite element computer program.
Fluid velocities of 10 ft/sec and 20 ft/sec were used in this evaluation. These velocities were equally distributed over a 14-inch length of tubing between the wrapper opening and the tube sheet. A fluid density of 50 lb/ft3 was used in the mathematic models.
4.5.6.1.2 ~Anal sis A tube is considered to be stable when the stability ratio is less than unity. The stability ratio (Ue/Uc) is defined as the ratio of the effective velocity (Ue) to the critical velocity (Uc). The effective velocity is a function of the distribution of secondary fluid flow velocity along the tube axis, fluid density, tube mass, and the mode shape of vibration. The critical velocity (Uc) is the threshold velo-city above which the tube amplitude increases as the secondary fluid velocity is increased and the induced energy from the fluid exceeds the damping energy dissipated by the tube. Computed fluid-elastic stability ratios for four tube configurations are p.esented in Figure 4.5.6-2 and 4.5.6-3 for the the fixed-fixed and fixed-pinned boundary conditions, respectively. Based on these results,.it can be seen that the undis-tressed cylindrical tube and the distressed flat tube bracketed the ten percent ovalized and kidney shaped configurations. The cylindrical tube was the more stable and the flat distressed tube the least stable.
Figure 4.5 .6-4 illustrates the marked increase in tube stability when the damping ratio for a fixed-pinned tube was increased from 0.01 to 0.025.
Presented in Table 4.5.6-2 are the computed bending mode fundamental frequencies for all tube cross sectional geometries. Since the cylin-drical and the flat are limiting configurations, analysis was extended to investigate the influence of an assumed 1000 lb compressive load being applied to these two tube configurations. Analysis was done for both the fixed-fixed and fixed-pinned, boundary conditions. Results indicated the fixed-pinned boundary condition to be the limiting condition for both tube geometries. The compressive load decreased the natural frequency of the cylindrical tube from 40.3 Hz to 38.5 Hz, and the natural frequency of the flat tube was reduced from 31.6 Hz to 27.9 Hz. In terms of fluid-elastic excitation, the reduction in natural frequency reduces the critical velocity and consequently reduces the stability margin. The reduced frequency also tends to increase the amplitude of vibration due to turbulence.
4.5-19
Cross-flow turbulence was evaluated, because it causes narrowband random vibration of tubes at about the natural frequency of tubes in the fluid. The vibration amplitudes vary randomly in time and direction.
Turbulence is thought to be the main cause of tube vibration in steam generators when the possibility of fluid-elastic excitation has been eliminated (Reference 4.5.6-1 and 4.5.6-2). Presented in Figures 4.5 .6-5 and 4.5 .6-6 are the computed amplitudes of vibration due to cross-flow turbulence as a function of cross-flow velocity for the cylindrical and flat section tubes. These amplitudes correspond to the natural frequencies presented in Table 4.5.6-2.
Axial flow turbulence can cause a tube to vibrate; however, cross-flow turbulence excites the tube. to larger amplitudes of vibration than does the axial flow condition. Amplitudes of vibration due to vortex shedding were not considered to be applicable, since vortex shedding is essentially a boundary layer phenomena and any condition that tends to disrupt the boundary layer will, in all probability, reduce the amplitude of vibration. Laboratory tests have shown no indications of resonance peaks due to vortex shedding in closely spaced tube arrays for the flow velocities considered relevant in the region of the wrapper opening and the tube sheet (Reference 4.5 .6-1). In most of the research investigations regarding vortex shedding, the flow velocity approaching the tube and/or an array of tubes has a relatively uniform velocity profile and low level of turbulence. However, in an operating >steam generator as the flow enters through the wrapper opening, the fluid flow becomes turbulent and the axial component of the velocity is thought to disrupt the boundary layer on the tube and the formation of vortices generated by the flow perpendicular to the tube. Although research continues to be done for vortex shedding in tube bundles, it may be considered a second order mechanism for inducing tube degradation in steam generators as the cross-flow impinges on only a small portion of the tube in the region of the tube sheet (Reference 4.5 .6-1). Of the three mechanisms identified with flow induced vibration, amplitudes generated by turbulence are smaller in magnitude than those generated by fluid-elastic excitation or vortex shedding. In closely spaced tube arrays, the considered predominant mechanisms are turbulence and fluid-elastic excitation.
4.5.6.1.3 Results In summary, the major results of this analysis are:
- 1. Cross-flow velocities in the range of 9 ft/sec can cause peak root mean square amplitude vibrations of .6 mils for a fixed-fixed cylin-drical cross section tube and approximately 1 mil for a fixed-pinned cylindrical cross section tube. The application of a 1000 lb compressive force had a negligible effect. Maximum amplitudes of vibrations would be roughly a factor of seven higher than the peak root mean square vibrations.
- 2. Cross-flow velocities in the range of 9 ft/sec can cause peak root mean square amplitude vibrations of 1 mil for a fixed-fixed flat cross section tube and approximately 3 mils for a fixed-pinned flat cross section tube. The application of a 1000 lb compressive force had a small effect. Maximum amplitude vibrations would be roughly -a factor of seven higher than the peak root mean square vibrations.
4.5-20
- 3. Fluid-elastic instability of the flat cross section tube for a fixed-fixed boundary condition will occur for fluid velocities in the range of 11.5-16 ft/sec. Since maximum cross-flow velocity is of the order of 9 ft/sec, fluid-elastic instability is not pedicted anal yti cal y. 1
- 4. Fluid-elastic instability of the flat cross section tube for a fixed-pinned boundary condition will occur for fluid velocities in the range of 10-14 ft/sec. Since the maximum flow velocity is of the order of 9 ft/sec, fluid-elastic instability is not predicted analytically.
4.5.6. 2 Local - Fl ui d. Ef fects 4.5.6.2.1 ~Per se A tear in a steam generator tube might result in a p otrusion which could be acted upon by fluid-induced alternating lift and drag loads to further increase the extent of the tear. The purpose of this section is to estimate the magnitudes of these loads so as to establish the role that fluid effects might have played in the shredding of tubes. The primary mechanism for p'oducing such loads is considered to be vortex-shedding.
4.5.6.2. 2 Anal ys i s Figure 4.5.6-7 illustrates the tear model used in the analysis. A protrusion with a height H and with a projected width D was exposed to a fluid cross-flow velocity of U. Alternating lift, drag, and torque acted as shown on the figure and were computed from the following equations.
LIFT = CL (pfU2/2) HD exp (i~t)
DRAG = CD (pfU2/2) HD exp (iut)
TORQUE = CL (pfU /2) (HD /8) exp (idiot) whe~e nf is fluid density, CL and CD are alternating lift and drag
~
coefficients, and ~ = 2~f is the circular frequency of vortex-shedding.
The latter can be computed from:
40 = 2 m f= 2 m S U/D where S = Strouhal number. From Reference 4.5.6-3 a Strouhal number of 0.2 and a lift coefficient of 0.5 were found to be approp.iate. From Reference 4.5.6-4, it was estimated that an alternating drag coefficient of approximately O.l times the alternating lift coefficient was app opr i ate.
4.5.6.2. 3 Results Figures 4.5.6-8 through 4.5.6-10 illustrate as a function of fluid velocity the variation of alternating lift, torque and drag loads acting on a tear whose dimensions are H = 2.5 inches and D = 0.5 inch. The 4.5-21
fluid velocity range covered is 0 20 ft/sec. This envelops the maximum between-tube cross-flow velocity which does not exceed 10 ft/sec (see Section 4 .5 .2).
Table 4.5.6-3 summarizes the alternating lift, drag and torque ampli-tudes and frequencies of excitation for a range of protrusion heights H and H/0 values of 1.0 and 5 .0. The fluid velocity is assumed to be 10.0 ft/sec In summary, the major conclusion from this analysis is that oscillating lift, drag and torque loads were not large enough to accelerate the tearing of tubing protrusions at the fluid velocities present in the tube bundle (< 10 ft/sec).
4.5.7 ~Fati ue This sub-section p esents the results of analytical calculations performed to assess the fatigue characteristics of steam generator tubing. Cases analyzed were:
a) a nominally plugged tube b) a plugged tube with a notch or stress riser type degradation due to foreign object impact c) a plugged tube with a locally collapsed section and notch or stress riser type degradation with and without continuous impact by a foreign object.
In the Case 3 analysis, lateral load magnitudes of 10, 25, and 50 lb were considered.
For the first two cases, no lateral impact loads were included and the fluid interaction loads were taken to be minimal, consistent with the results of sub-section 4.5.6. Consequently, Cases 1 and 2 were examined from the viewpoint of low cycle fatigue due to thermal-mechanical operating loads and cycles. On the other hand, Case 3 analysis considered fluid-interaction and cyclic lateral impact loads and as such is of primary interest from the viewpoint of high-cycle fatigue. For
'ase 3, worst case axial loads were assumed in conjunction with the fluid interaction analysis.
In addition to the gross fluid interaction evaluations discussed above, the case of a locally shredded tube is considered in sub-section 4.5.7.2 to evaluate the propensity of crack propagation and continued shredding under the influence of local fluid effects.
The following nomenclature is used in the remainder of this section.
Pp Primary side pressure, psi Ps Secondary side pressure, psi Inside radius of tube = 0.3925 in ro Outside radius of tube = 0.4375 in 4.5-22
ET = Nodulus of Elasticity for tube at operating temperature, ps 1 Do = Ou ts i de di ameter of tube = 0.875 in Di Inside diameter of tube = 0.785 in L = Length of tube span = 52.2 in e = Tube sheet rotation angle = 0.096 TTP Mean temperature of plugged tube, F TTA = Mean temperature of active tube, F TS8 = Mean temperature of stub barrel, 'F a = Coefficient of thermal expansion, in/in-F I = Moment of inertia for tube, in4 Displacement, in or displacement per unit length, in/in 4.5.7.1 Tubin Evaluation The following parameters are assumed for the fatigue evaluation of a plugged tube throughout this section.
Design minimum tube wall = 0.045 in Maximum ovality = two percent for straight length tubing Stress concentration due to notch or stress riser type degradation = 4.0 Initial TS and TSP holes offset = 0.1 in Additional-ly, the evaluation included the worst case axial loads based on, hot leg parameters. Tubes on the cold leg and/or away from the wedge areas would have significantly la~ger fatigue margins than that indi-cated from the analysis due to the increase in tube suport plate flexibility away from the wedge areas.
Case 1:. Nominally Plu ed. Tube The stress categories used in calculating the stresses are listed below. The calculated stresses are tabulated in Table 4.5.7-1. The sunmation of the principal stresses is given in Table 4.5.7-2. The alternating stress and the usage factors are given in Table 4.5.7-3.
Stress Cate pries a) Pressure stress in the tube:
4.5-23
~radial b) Bending stress in the tube due to rotation of tube sheet as a result of p imary-to-secondary zP:
3ETD bending c) Bending stress due to in-plane thermal growth mismatch between tube sheet and support plate:
3ETD 6TH bending L2 where:
6TH = Radial expansion differential between tube sheet and tube support plate.
L(c aT)TS (~ aT)TSPj R and R 56.5 in, the distance of the outermost tube from the steam generator centerline.
d) Bending stress in the tube due to as-built offset tto 3ETD t bending L2 e) Axial stress in the tube due to thermal growth mismatch between tube and stub barrel.
axial = ET ((c '>T)SB (e ~T)TP]
f) Bending stress in the tube due to axial forces acting through tube offset, a.
"A' o bending 2I where:
~
"A = (g ~axial) . (Area of tube)
= aTH + ITS + ~0 ROT.
4 4
~ (D, D,.)
I = = 0.01033 4.5-24
g) Axial stress in the tube due to thermal growth mismatch between plugged tube and active tubes.
~axial = ET f(a hT)TA (n hT)TP]
h) Hoop bending stress due to two percent ovality. This was obtained from the results of finite element analysis in Section 4.5.5.
Case.2:- Nominally Plu ed Tube with a Notch This case analysis is basically the same as the Case l analysis, except that a stress concentration factor of 4.0 was applied to the axial stresses to increase the stresses due to a notch or stress riser. The sumaation of principal stresses and the usage factor for this case are given in Table 5.3.7-4.
'I In both of the above cases, the hot standby feedwater cycling (following the plant heatup but p ior to plant unloading) was analyzed for a maxi-mum aT on the hot leg side. During a 300 GPM slug feed cycle, the maximum aT between the plugged tube and the stub barrel would be 71 F.
Use of this wT in the case of the Ginna Steam Generator is conservative since the feedwater transient at Ginna since early l980 is a 35 GPM continuous feed type cycle, which would result in a lower sT than a slug feed type cycle. The number of feedwater cycles, consistent with the plant heatup cycles, was assumed to be 200 over the 40 year design life of the plant.
Case 3: -P lu ed. Tube. with Locally-Collapsed. Section This section presents the results of the evaluation of a plugged tube with a locally collapsed section and notch or stress riser type degra-dation with and without continuous impact by a foreign object. For the analysis, the tube was assumed to be locally collapsed over a 2.0-inch length beginning at 4.0 inches above the tube sheet. An, analysis, as described in Section 4.5.6 was performed to obtain the frequency and the root mean square stresses, due to cross-flow turbulence. A 1000 lb axial load on the tube was considered in this analysis to simulate the effect of thermal growth mismatch loads due to axial restraint at the first tube support 'plate.
The uncollapsed portion of the tube was assumed to have a nominal two percent ovality. Hoop stresses on oval tubes under the combined effects of external p essure and various magnitudes of lateral loads were obtained from the detailed finite element analysis described in Section 4.5.5.
for a fixed-pinned condition, the maximum peak stress*, due to cross-flow turbulence with a fluid velocity of 10.0 ft/sec and including a
- The analytically calculated stresses correspond to a root mean square (RMS) excitation. For an actual flow field in a typical steam generator, cross-flow turbulence causes narrow-band random vibration of tubes. Based on experimental data, the ratio of peak to RMS stress is approximately 7.0 and is used in this evaluation.
- 4. 5-25
stress concentration factor of 4.0 for a notch, was equal to + 11.24 ksi which is less than the material endurance limit of approximately 13 ksi at 1011 cycles. Hoop stress due to two percent ovality and various impact loads was:
0 lb impact = 15.0 ksi 10 lb " = 19.5 ksi 25 lb " = 26.0 ksi 50 lb " = 39.0 ksi Using a high-cycle fatigue curve, the all'owable cycles for the above loads are as summarized in Tables 4.5 .7-5 and 4.5 .7-6. This high-cycle fatigue curve is proposed to ASME to include the alternating stresses representing 1011 cycles for Inconel material. The times to failure (cumulative usage factor = 1.0} in Table 4.5 .7-6 are based on an impact frequency of 38.5 Hz (3.3 x 10> cycles/day) which is the fundamental vibration frequency of a fixed-pinned tube with a 1000 lb axial compressive load.
Based on results of the fatigue analysis above, both nominal and degraded plugged tubes without a lateral impact have acceptable margin to fatigue due to the worst case thermal-mechanical and hydraulic loadings over the design life of the plant. A plugged tube under axial compression, with a locally collapsed section and a notch or stress riser, subjected to a continual lateral impacting load will fail in fatigue; the calculated failure time for this case varies from about a day to a few weeks depending on the magnitude of the impact loading.
4.5.7.2 Shred. Evaluation A tear in a tube could result in a p.otrusion which could be acted upon by fluid-induced alternating lift and drag loads to further increase the extent of the tear. For this case, the hydraulic load calculations were discussed in the previous section (4.5 .6). Even for significantly large tears (0.5 inch wide x 2.5 inches high), the calculated loads were very small, resulting in maximum tube stresses well below 10.0 ksi. Thus based on a high-cycle endurance limit of app.oximately 13 ksi at 101i cycles as previously given, no crack propagation is expected.
4.5.8 Wear The purpose of this section is to calculate the period of time for sufficient wear to occur that bursting of a tube woul'd result. The calculation will consider potential ranges of wear for Inconel tubes rubbing agains t other I neon el tubes.
Oewees (Reference 4.5.8-1) previously gave the value of 582 in2/lb ys the specific wear coefficient for Inconel sliding on Inconel in 500 F water and 8540 in2/lb in 300 F water. He also P.esented data f'r other material combinations and deduced that fretting wear is about 1/10 that of continuous sliding wear. This would mean that the range for fretting wear of Inconel on Inconel should be 58 in2/lb to 854
'in2/lb or an average of 456 in2/lb.
4.5-26
4.5.8.1 Wear-Cal cul ations The emperical wear formula according to Archard's theory on wear is given as:
V = CFT x 10-12 where: V = wear volume, in3 C = specific wear coefficient, (psi)-1 or in2/lb F = force, lb T = total travel or rubbing distance, in It can be shown that a uniformly-thinned 7/8" 0.0. tube will burst when the wall is reduced to app oximately 6.6 mils under the normal Ginna operating conditions, assuming an ultimate strength of 89 .7 ksi .
(See Section 4.5.8.3.)
Therefore, if the rub area is 8 .00 inches x 0.10 inch and the wear depth is 0.0434 inch which is reasonably similar to the Ginna situation, V = 8 x 0.1 x 0.0434 = 0.0349 in3.
For purposes of establishing a contact force for one tube rubbing on another, it is assumed that a severed tube is catilevered from the tube support plate and is 50 inches long. Under fluid drag forces, the severed tube will lean against a nei ghboring tube that is supported both at the tube sheet and at the tube support plate.
The drag force, Fp, is computed as follows:
Fp = 1/2 p AU2 Cp where: p density of water = 1.94 slugs/ft3 0.001123 slugs/ in3 projected area of the severed tube in the wrapper opening 12 x 7/8 = 10.5 in2 flow velocity = 3 ft/sec 36 in/sec dr ag. coefficient 0.8 (assumed turbulent)
Fp 1/2 x 0.001123 x 10.5 x 362 x 0.8 = 6.11 lb Under this load, the severed tube deflection is 0.5548 inch.
=~
Since the neighboring tube is 0.31 inch away, the deflection of the tube is limited to this clearance. The contact force F can be estimated as:
F 6.11 = 2.7 A value of F = 3 lb will be used.
4.5-27
With C T =
= 456 CF 8; 8349 (psi)-1
-.Y- -
10-r2 x.le
=
~
and F = 3 V-x 12 lb,
= 25.5117 x 10 6
in When the severed tube is leaning against a neighboring tube, it will be assumed that the neighboring tube will vibrate at its first natural frequency while the severed tube maintains contact with it. Since the first bending frequency of the neighboring tube, which is roughly 52 inches long with fixed/partial restraint end conditions, is about 50 Hz, this frequency will be taken as the rubbing frequency between the two tubes. Based on the observed wear widths of 0 to 100 mils (circumfer-entially) on the tubes removed from Ginna, the mean amplitude of motion will be taken as 50 mils.
Defining T's the rub velocity, then T'=2xex f
= 2 x 0.050 x 50 = 5 in/sec If "t" is the total time to rub the tube thickness down to the minimum wall before burst, T- 25;5117.x 18.6 5 lQ23 x 1Q6 T'
= 59 days For an upper bound, "t" could be longer in proportion to the specific wear coefficient, or t = 59 x 456/58, = 1.27 years.
. The time periods calculated above reasonably envelope the observed wear range. For example, burst tube R42 C55 went from less than 40 percent to 84 percent wall thinning between May 1981 and January 25, 1982.
4.5.8. 2 Wear by Severed. Tube The wear patterns seen in the Ginna B-Steam Generator point to rubbing the Q.D. of a neighboring tube by a severed tube. In particular, the cut tube would lay against the next tube and then would fret trans-versely against it as a result of fluid-induced motion. In so doing, the severed tube would wear down the walls of the adjacent tube(s).
In contrast, an unsevered tube would not be able to inflict such wear damage on surrounding tubes for lack of contact. Mechanical, thermal
- 4. 5-28
and hydraulic loads are insufficient to cause tube deflection sufficient for tubes restrained at both ends to interact, as evidenced by the lack of tube wear in regions mid-way between the tube sheet and first support plate. That is, wear by unsevered tube(s) does not represent the mechanism seen at Ginna.
4.5.8.3 Burst Stren th of-Tubes The purpose of this section is to determine the minimum wall thickness to preclude bursting of an active tube and to correlate the analytical prediction with the actual wall measurements of the burst tube (R42 C55) from the Ginna 8-Steam Generator.
The tube wall thickness for burst during normal operation is computed by assuming uniform thinning and using the pressure stress equation in NB-3324.1 of the ASME Code. In this case, however, the primary membrane stress intensity used will be the ultimate strength of the tube material.
For the Ginna tube:
Pm 89.7 ksi Ri = 0.397 in P ~
= 2250 psia Po = 780 psia
= 2250 780 = 1470 p i
.-1479-x 0-397 ----.
t=
= 0.0066 in Measurements of the burst tube from Ginna showed, that the thinnest wall was 0.008 inch. Thus, the thickness based on a rather simplistic calculation correlates well with the actual measured thickness.
4;5-29
References for Section.4;5 4.5.2-1 A. C. Spencer, "Method of Characteristics for Solving Two-Dimensional Reactor Core Flows", proceedings of the 1973 Conference on Mathematical Models and Computational Techniques for Analysis of Nuclear Systems", CONF-730414-P1(ANS), 1973, pp. III-3 to III-22.
4.5.2-2 WECAN: Westinghouse Electric Corporation ANalysis, User's Manual, Second Edition, March 1981.
4.5.4-1 J. C. Hunsaker and R. B. Rightmire, "Engineering Applications of Fluid Mechanics," McGraw-Hill Book Co., 1947, pp. 198-204.
4.5.6-1 M. J. Pettigrew and D. J. Gorman, "Experimental Studies in Flow-Induced Vibration to Support Steam Generator Design, Part III. Vibration of Small Tube Bundles in Liquid and Two Phase Cross Flow", AECL 5804, June 1977.
4.5.6-2 H. J. Connors, "Flow-Induced Vibration and Wear of Steam Gen-erators", Nuclear Technology Volume 55, November, 1981.
4.5.6-3 R. E. D. Bishop and A. Y. Hassan, "The Lift and Drag Forces on a Circular Cylinder Oscillating in a Flowing Fluid", Proceed-ings of the Royal Society of London, Series A, Vol. 277, pp.
32-75, 1964.
4.5.6-4 M. J. Sox, "Fluctuating Pressure Field on a Yawed Rigid Cylin-der", Ph. D. Thesis, University of Minnesota, December 1978.
4.5.8-1 DeWees, N. B., "Sliding-Wear Performance of 135 Material Couples in High Parity Water at 300 F and 500 F," Bettis Atomic Power Laboratory, WAPD-314 AEC Research and Development Report, Control AT-11-1-GEN-14, May 1967, Pittsburgh, PA.
4.5-30
TABLE 4.5.2-1 NOMINAL AND BETWEEN-TUBE CROSSFLOW VELOCITIES IN AND NEAR THE PLUGGED TUBE REGION Cross flow V el oci ty
--- (nominal)---.
Crossflow Velocity Location one cell-of. tubes. lu ed Perimeter Cell 8.2 ft/sec 8.91 ft/sec (Face A)
Per imeter Cell 8.2 ft/sec 8.91 ft/ sec (Face C)
Perimeter Cell 9.01 ft/ sec 9.11 ft/sec (Face B)
One Cell in from 8.54 ft/sec 8.84 ft/sec P erimeter (Face B)
Two Cells in from 8.21 ft/sec 8.56 ft/sec P erimeter (Face B) 4,5-3l
TABLE 4.5.3-1
SUMMARY
OF WORST CASE AXIAL LOADS ON A,PLUGGED TUBE Source Type (no;of cycles) lb."
Max. Value
.. Analysis
~A1 l1 r
- 1. Secondary Pressure Constant -604. F 1 ow Induced (maximum at no-load) (compression) Vibration
- 2. Tube-to-Shell Interaction
- a. Hot standby with Transient (200) -1940. Fatigue 70 F feed
- b. Plant loading (14,500) 825. Fatigue
'ransient
- c. Plant unloading Transient (14,500) -1375. Fatigue
- d. Full power Constant -550. Flow Induced steady-state Vibration
- 3. Plugged-to-Active Constant 1410 Col 1 apse Tube Interaction (Full power steady-state)
- Pressure and thermal interaction loads cannot both occur on the same tube. However, tube-to-shell and plugged-to-active tube interaction loads can occur simultaneously on a tube. The actual load magnitude on a plugged tube would depend on the location of the tube and the number of active tubes around it. Assuming superposition of steady-state loads (Cases 2d and 3), tensile loads of the order of 1000 lbs can be expected in a plugged tube near a hot leg wedge location. Axial loads on tubing in cold-leg and/or regions away from the wedges would be lower due to smaller temperature differentials and support plate flexibility.
0.5-32
TABLE 4.5.4-1 FOREIGN OBJECT INDUCEO LOAOS Model 1* Model 2*
Xp (0) Peak Load Peak Load Case ~inches Uss (ft/sec) UA (ft/sec) lbe -- ~lbs 2.3 26 107
.69 .69 .15 .28
.345 .69 .39 1.40
- Model 1: K5G = 3717 lb/inches MSG = .03767 lb sec2/inches Model 2: KSG = 19439 lb/inches MgG = .01068 lb sec2/inches 4.5-33
TABLE 4.5.6-1 ANALYSIS PARAMETERS Outside diameter 0.875 inch Inside diameter 0.775 inch Tube pitch 1.234 inches Flexural moment of inertia O.ill inch4 Axial area 0.1290 inch2 Tube density 0.304 lb/inch3 Modulus of elasticity 29.2 x 1061b/inches2 Secondary fluid density 50 lb/ft3 4.5-34
TABLE 4.5.6-2 TUBE FUNDAMENTAL FREQUENCIES FOR THE VARIOUS CASES ANALYZED Cross Section Fixed-Fixed Fixed-Pinned P=O P=O P= -1000 lbs Cylinder 58.7 HZ . 40.3 HZ 38.5 HZ Flat 48.9 HZ 31.6 HZ 27.9 HZ 10 Oval ized 58.3 HZ 40.0 HZ Kidney 58.2 HZ 39.9 HZ 4,5-35
0 1
TABLE 4.5.6-3 VORTEX-SHEDDING INDUCED LOADS ON PROTRUSIONS FOR A FLUI.D.
VELOCITY OF 10.0 FT/SEC Lift Torque Lift a Torque Drag Drag H/D H. Inches ebs in ebs Frequency HZ) (ebs) Frequency Hz 1.0 O.l .28 (10-2) .35 (10 4) 240 .28 (10 3) 480 1.0 0.20 .11 (10-1) .28'(10 3) 120 .11 (10 2) 240 1.0 0.30 .25 (10-1) .95 (10-3) 80 .25 (10-2) 160 1.0 0.40 .45 (10-1) .22 (10-2) 60 .45 (10-2) 120 1.0 0.50 70 (10-1) .44 (10-2) 48 .70 (10-2) 96 5.0 0.50 .14 (10-1) .18 (10-3) 240 .14 (10-2) 480 5.0 1.00 .56 (10-1) .14 (10 2) 120 .56 (10 2) 240 5.0 1.50 0.13 .47 (10-2) 80 .13 (10-1) 160 5.0 2.00 0.22 .11 (10-1) 60 .22 (10-1) 120 5.0 2.50 0.35 .22 (10-1) .35 (10 ) 96 4,5-36
TABLE 4.5.7-1 CALCULATED STRESSES (KSI)
NOMINALLY PLUGGED TUBE a a a b c d e g h Transient. ~A ~H ~R ~B ~B ~B ~A ~A ~BH Plant Heat Up 0.0 -9 .77 -0.50 2.42 0.11 2.81 -4.96 3.02 0.0 22.0 P lant Cool down 0.0 -9.77 -0.50 2.42 0.11 2.81 -4.96 3.02 0.0 22.0 Plant Loading 0.0 -7.74 " -0.40 2.42 0.23 2.81 +4.54 2.85 -+3.36 17.2 P lant Unloading 0.0 -7.74 -0.40 2.42 0.23 2.81 +4.54 2.85 +3.36 17.2 Step Load + 10 0.0 -7.74 -0.40 2.42 0.23 2.81 +4.54 2.85 +3.36 17.2 Reactor Tri p 0.0 -7.74 -0.40 2.42 0.23 2.81 +4.54 2.85 +3.36 17.2 50 Step Load 0.0 -7.74 -0.40 2.42 0.23 2.81 +4.54 2.85 +3.36 17.2 Decrease Hot Standby 0.0 -9.97 -0.50 2.42 0.11 2.81 +3.21 6.04 +9.94 22.0
- Largest axia stress is used to calculate this bending stress.
TABLE 4.5.7-2 PRINCIPAL STRESSES NOMINALLY PLUGGED TUBE (KSI)
Transient Ax i al- Hoop Radial Plant Heatup -13.32 -31.77 -0.50 Plant Cool down -13.32 -31.77 -0.50 Plant Loading 12.85 -24.94 -0.40 P 1 ant Unloading 12.85 -24.94 -0.40 Step Load + 10 12.85 -24.94 -0.40 Reactor Tri p 12.85 -24.94 -0.40 50 Step Load 12.85 -24.94 -0.40 Increase Hot Standby 21.32 -31.77 -0.50 4.5-38
TABLE 4.5.7-3 USAGE FACTOR FOR A NOMINALLY PLUGGED TUBE Al 1 owab1 e Usage Transient Cycles Sa*,ksi Cycles- Factor Plant Heatup 200 20.4 106 0.0002 P lant Cool down 200 20.4 106 0.0002 Loading 'Plant 14,500 25.6 106 0.0145 P lant Unloading 14%500 25.6 ]06 0.0145 Step Load+ 10 4,000 25.6 106 0.004 Reactor Tri p 400 25.6 106 0.0004 50 Step Load 200 25.6 106 0.0002 Decrease Hot Standby 200 36.9 105 6: 662.
TOTAL 0.036
- Sa is adjusted to E = 26 x 10 ksi.
4.5-39
~ ~
USAGE FACTOR FOR A NOMINALLY GED TUBE WITH NOTCH OR STRESS RISER
-... Princi pal. stresses;- ksi- Al 1 owab 1 e Transient xl a a Hoop Raia. Cycle Sa*,ksi -Cycles-- Factor Plant Heatup -39.96 -31.77 -0.50 200 29.0 300,000 0.0007 Plant Cooldown -39.96 -31.77 -0.50 200 29.0 300,000 0.0007 Plant Loading 38.55 -24.94 -0.40 14,500 42.0 50,000 0.29 lant Unloading 38.55 -24.94 -0.40 14,500 42. 0 0.29 P
50,000'sage 50,000 Step Load + 10 38.55 -24.94'0.40 4,000 42.0 50,000 0.08 Reactor Trip 38.55 -24.94 -0.40 400 42.0 50,000 0.008 50 Step Load Decrease 38.55 -24.94 -0.40 200 '2.0 0.004 Hot Standby 63.96 -31.77 -0.50 200 39.6 70,000 0;003 TOTAL = 0.6764 Stress concentration factor of 4.0 is applied for the notch or stress riser.
- Sa is adjusted to E = 26 x 103 KSI
TABLE 4.5.7-5 STRESS AMPLITUDE AND ALLOWABLE CYCLES WITH LATERAL IMPACT LOAD Im act.k.oad .ksi ~S*,ksi No; of C ches Allowed 0.0 9.22 10.0 14,15 108 25.0 16.81 8.5 x 106 50.0 22.14 2.4 x 106
- Sa is adjusted to E = 26 x 103 ksi 4.5-41
TABLE 4.5 .7-6 POSTULATED FAILURE TINE WITH LATERAL IMPACT LOAD Impact Loads Al 1 owab le Postulated Failure
~
ebs) -- ~
Cycl es/Day - Cycles- ~
Time**(days 0.0 3.33 x 106 10.0 3.33 x 106 108 30 25.0 3.33.x 106 8.5 x 106 2.5 50.0 3.33 x 106 2.4 x 106 0.7
- Failure is defined analytically when cumulative usage factor = 1.0.
4.5-42
WRAPPER
.045" OUTER PERIPHERY TUBE,.875 O.D.x 055lltt a 1 0.75 SUPPORT BLOCK WEDGE PLATE SUPPORT UJ Lal 5T,.S I STUB BARREL/SHELL
- 14. 00'6. 5" 22.00" TUBE SHEET ATIC SHOWING HOT LEG FIGURE 4 5.1 HE SPAN GEOMETRY
- 4. 5-43
Dst ESP.
IDES
'8 27 26 25 24
'3 22 RAPPER~=-
20 19 18'7, 16 15 14 I 13l 12 10 7,
2;14 ft/se<
51 .
u( rNLET) =
2/s ec. 3
+3.45 I'e/sec Y J I
3 19 20'2"I
,EX-5 85'I TUBE SHEET
-'IGURE 4 5.2-1 CHARM MODEL FOR RGE
6.
erst Hpo r Sup Ii M% ih']
g!Q t I) z 1
h;
)
i/!
t ti g3 ~ ending Zu>e Sheet" I I
i t~
l Wrap per g I
i I t I I FIGURE 4,5,.2-2 THREE-DIhKNSIONAL (WECAN) ifiYDRAULIC
- hX)DEL OF TUBE SHEPZ TO FIRST SUPPORT PZATE REGION c
8 73 I
1 ~ 9 16 4 02 109 04 105 77 79 27 29 ES PLUGGED 5 TUBE P TTCHES IGURE 4.5.2<<3 YDRAUL mz)
- 4. 5-
e 7
l5 21
<4 I
6 15 I
16 h.2 ubes 20'ugged P
206 JQ +
201~
FIG URE 4, 6.2 '4 ST TION ODD S F OP, RAULIC SIP (Z 7 I
p 5-47
AVERAGE VELOCITIES (LATERAL AND AXIAL) 0 1.92 ~ 74 ~ 93 .89
- 2. 15 1. 71 1.22 .93 3.94 2.57 2.08 1. 70 1.37 1.03 .66 0 2.16 1.70 1.17 .84 .76 1.55 3.94 3.51 2.61 1. 98 1.53 1.15 .76 0 1.84 1.48 1.00 .71 ~ 63 1.28 3.94 3.53 2.66 2.02 1.56 1. 19 .82 0 1.44 1.18 ~ 81 .56 .48 1.02 2.38 ft/sec 3.94 3.56 2 '1 F 05 1.60 1.24 .88 0 1.06 .89 ~ 61 .41 .35 .76
- 3. 94 3. 58 2 ~ 74 2.08 1.63 1.28 .94
.70 ~ 59 .41 ~ .27 ~ 23 .50 3.94 3 59 2.76 2. 11 1.66 1.33 1.00
.36 ~ 30 ~ 20 .14 .25 3.94 3.60 2.78 2.13 1.69 1.37 1.06 Lataral Valocityt ft/sec +
Axia1 V locity, ft sec +
X~ 1 FIGURE 4.5.2-5: GINNA NOMINAL CASE 4.5-48
NO HEAT LOADS ONLY
.19 ~ 84 1.01 .99 .86 1.91 1.73 1.27 .95 ~ 78 3.94 2 '9 2. 19 1. 80 1.46 1.15 ~ 85 1.91 1.71 1.21 .87 .70 1.05
- 3. 94 3.57 2.74 2.09 1.64 1.27 .95 1.64 1.48 1.04 .74 ~ 58 .81 2.38 ft/sec 3.94 3.59 2.78 2.13 1.66 1.30 .99 1.28 1.18 .84 ~ 58 ~ 45 .57
- 3. 94 3. 61 2.81 2.16 1.68 1.33 1.04
.95 .89 .63 ~ 44 .33 .35
- 3. 94 3.62 2 ~ 84 2. 18 1.72 1.37 1.08
.63 ~ 59 .42 .29 .22 .19 3.94 3. 63 2.86 2.20 1.74 1.40 1.13 ll 3.'94
.32 3.64
,.30 F 88 2'6
.21
>22, 0
.14
.'.3..
0
~
1.43
.09 1.17 0
Lateral Velocity,
+
Axia3. V locity,
't/sec ft sec +
X~ 1
~-
FIGURE .4 5.2-6: GINNA WITH TUBES PLUGGED
- 4. 5-49
1st TSP l
i 1
i i'll l) 1 1 I I i I t r e
~
~
~
~
~ ~ 0 ~
ti))l l I i e-~
\
0 0 0
0 0
0 ~
))i)tl
) ) I i))1 l'1 1 t I
r r s 0
0 0
e 0
e 0
0
~
~
>>) ~ti
)
i 11 t
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a a
e e
e 0
0 e
0
~
a
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~
I 1:~ a e 0 S I I l 1 i a a 0 . a r i
r I I I
I 1>>i 1 1 a 1.5 a
ft/sec.
r I 1'l) 1 1 00
'i I r r I
\ I t 1 i a 1
t
~ e 2.38 ft/sec.
2.14 ft/sec.
I
~ 4 a 0 0 a 40T SIDE COLD SIDE FIGURE 4 5 2-7 GINNA BASE CASE
- 4. 5-50
0 REGION OF REGION OF TUBES PLUGGED TUBES PLUGGED
's
~ ~
I. r r ~
ZONE OF ~ o ~ ~ ~ ~
ZONE OF DOWNWARD I ~ t ~ ~ o ~ ~
DOWNWARD VELOCITY I I ~ ~ ~ 0 S ~ ~
VELOCITY ft/sec MAX~.62 I ~ ~
0 0
'Is 0
o 0
~, a-,'Qyy ft/sec gs 4 O ~ ~ ~
\ % a o e o o ~ ~
1 e e e s l ~
1l 1 l 1 1 a i I I
> 1 U < 1.5 ft/spc kg j
~ ~ I i 1 l i
- i. r r I i l,l i1'li i I I I I r r.
r ( l r r r I 2.38 ft/sec r 2.14 ft/sec a e r ~
I ss s I HOT SIDE COLD SIDE
'I FIGURE 4 5,2-8 GINNA WITH TUBES PLUGGED 4.5-.51
0 HOT SIDE COLD SIDE FIGURE 4g5 2-9 QUALITY DISTRIDUTIOtl OF GINNA K'-BE CASE
1st TSP" r
Vl I
Vl GD TUBE SHEET HOT SIDE COLD SIDE FIGURE 4.5.2-10 QUALITY DISTRIBUTION OF GINNA WITH TUBES PLUGGED
~
~
I I I
l I I f
~ I i I
I I t
~ I I I I
I 1 I
IC ~i 0 N U I I
~ I I
GORE 4. 5.2 EMCH'IES Ce I
I
. 5-54 J~
0 ap er be U
UBE B NDL E REG'ION WI HFO U
P re 4 .5.4 YNAMIC MODEL 0 FOIE IGN'BJ ECT IMPACT
- 4. 5-55
- 30. 0
- 20. 0 Model 1~
LU CQ Case 1 0.0
+P CA
-10. 0
-20.0
.1 2 3 .4 5 6 7 8 9 10 TIME(SECS)
SG TUBE FORCE"VS. TIME Figure 3'> 4"2 " ', SG TUBE SPRINB~E -TIME HISTORY QTEADY-STATE FLUID VELOCITY = 2.3 ft/sec ALTERNATING VELOCITY ~,0),,..
1.0 0.0 5-1;0 Model 1~
Case 1 o-2 0 C/lI fop,g.i 9 n 0 Je
-3.0
-4.0
-0.0 /
2 3 M -5 '6 7 8 9 10 TI (SECS}
P OSITI N VS. T IME
. Fiaure -
M~-3 .... '5'QREIGN~~@$ 5G TUBE DISPLACQEN
".TIHE 'HISTORY
'STEADY-STATE FLUID VELOCITY = 2.3 ft/sec ALTERNATING VELOCITY =0)
~-56
0.30 0.20 lA Pl Nodel 1~
O 0 0.10 Case 2 0.00
-0.10 10 20 30 40 50 60 70 80 90 100 TINE(SECS)
S jG TUBE FORCE VS. TIME Figure 4' 4."4. S/G TUBE SPRING FORCE-TINE HISTORY (STEADY STATE FLUID, VELOCITY ~ .69 ft/sec. ALTERNATMG VELOCITY=
.69 ft/sec.
7.PE-05 6.0E-05 TU M 5. OE-05
- 4. OE-05 H Hodel 1~
R Case 2 o 3.05-05 M
Foreign Obje t 1.OE-05 TO" 20 30 40 50': 60 70 80 90-'i 100
'IME(SECS)
'tOSITQN'VS. TINE Figure '4.5.4-5..':,; .:: ..'lREIGFIFLUID VELOCITYR =
UBOECT AND TUBE DISPLACENENT (STEADY STATE .69 FT/SEC.
ALTERNATING VELOCITY= .69 FT/SEC.
4.5-57
- 0. 50 0.40 0-30 HodeL 1~
.0. 20 l Case 3 0
N
'0.1 0.00
-0. 10
-0.20 10 20 30 40 50 60 70 80 90 100 TINE(SECS)
S/G TUBE FORCE VS ~ T3:ME figure:;4.'. 5'4 .6 S.G. TUBE SPRING FORCE = TINE HISTORY (STEADY'STATE FLUID VELOCITY = '.345 FT/SEC ALTERNATING VElOCITY =.69 FT/SEC ) ~ ~
~, 0.001 uh
~ ~
- 0. 000 Cxl O
aH tlodel 1~
O a Case 3
-0. 001 H
M r Q
C4
-0. 002 Foreleg n Objec
-0.003 Q,33 ~ 'l0j 20 j30 '. 40
', 5100
'TPK(SECS)
POSlQ5M:VS. TINE FOREIGN GNKCX4ti SG TUBE DISPLACQgKE=TIflE HISTORY (STEADY'STATE FLUID VELOCITY = .345 FT/SEC ALTERNATING VELOCITY = '.69 FT/SEC )
- 4. 5-58
s IICIIII IL ss los'4 Is I ss ooooy o~ s, s ~
~
- I
~ Tl I I't I IT I
~
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II:II II': I
~
O.D. x 0.050" THICK TUBE
~ ~
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G OF A II ~
NA RIIU R R 14L L'0
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o C E
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~ ~ ~ ~ I I FIGURE 4.Q.Q-],: EXTERNAL PRESSURE AND AXIAL LOAD REQUIRED FOR INCIPIENT YIELDING
b'(uopgcJ I= o kora a o X (dwDiaz) N4C <c. Cooed. SYm~
IX%
Co pg a AlopE. Afo.
a SC,Scour. yc.
C p (ev~c) n PA6fs'vZE ZoADnte ou OwER S'c/MFA c'E.
CFAca g,)
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se P
S 4
3 a"
~$
~~(Z) g a4 VV>gO7K a O Afapes 4 rug@ 4S: d'p <<a kc&c g4FX wp ppgg gpyy' y ALL ELEMENTS ANSYS STIF48 COMPUTER MODEL FOR UNIFORM PRESSURE LOADING FIGURE 4.,5. 5-2 4.5-6;0
.O7O" (TYPl
.035" ALL ELEMENTS ANSYS STIF63 MODEL USED FOR CONCENTRATED LOADING FIGURE 4.5.5-3
4,5-6f
GLOBAL Y UY ~ ROTZ ~ 0 UZ ~ ROTY ~ 0 Y (TANGENTIAL)
NODAL COORD'.
X (RADIAL)
.070 TYP.
UY~ ROTY~ 0 10'YP.
GLOBAL X ELEMENT NO. 2 ELEMENT ND. I
~
GLOBAL Z
.O3S UY ~ ROTZ 0+
ELEMENT NO'S. 1 5 2 LOADED WITH INWARD ACTING RADIAL PRESSURE BOUNDARY DISPLACEMENTS IN NODAL COORDINATE SYSTEM BOUNDARY CONDITIONS FOR CONCENTRATED LOAD MODEL FIGURE '.5;5-4
- 4,5-%
480
.06 440 .04
~ ~ ~ ~ 4 I ~
.02 400 ..02
-1000 LBS e = .01
=-... TENSILE LOAD:-'-", ..=w ~:"
~ :~ .
~ '.-
360 320
~ ~ ~ ~~1
~ ~
~
~
~ 'a
~ ~
~
280 ~~ ~
240
~ ~
E 200 I 160 0;875 in O.D. x .030 in WALL:;"
120 80 MAX. YIELD STRENGTH = 57.7 KSI ....;::
40 0
':=.-. : .'IN.
~
YIELD STRENGTH = 35.5 KSI (AT 550'F) ~ ~ l
.06 200 8.04
~ ~ w ~
160 .02 e = .Ol 120 1000 LBS TENSILE LOAD i 80
- -~ ~
40
.-.l:0.875 in
- =. O.D. x .050 in WALL .-::-:
- MIN. YIELD STRENGTH = 35.5 KSI -.-"-
0
.100 150 . 200 250 300 CONCENTRATED RADIAL LOAD - (LBS)
Figure 4.5 5-'5.::..
MAXIMUM STRESS IN TUBES OF VARIOUS OVALITIES LOADED WITH 1000 PSI UNIFORM EXTERNAL PRESSURE PLUS CONCENTRATED RADIAL LOAD 4.5-65
.030 in WALL STRESS (KSI) 2
.050 in WALL 0
0 .2 3 ~ 4 AXIAL DISTANCE FROM LOAD (IN)
HOOP STRESS Itt OUTER FISER DUE TO 5 POUND RADIAL LOAD FIGURE - 4.5.5-6 4.5-g:
I
.875 IN O.D. .05 IN NOM. WALL LINEAR DIMENSIONS "-IN VELOCIT Y= FPS YI SECONDARY FLUID CROSS FLOW REGION FLUID VELOCITY 14.0" 6.0" II p
I I I FIGURE 4
4.5,8-l;:BASIC ANALYSIS MODEL GEOMETRY'
FIXED-FIXED BOUNDARY CONDITIONS DAMPING RATIO = 0.01 e
SYIISOL CROSS SECTION 1ST MODE FRE . - HZ
~
0 CIRCULAR 68.7 I., FLAT .,".:;--:48;9 0 10K OVALIZED :.:.: .58;3 A
KIDNEY -':":. <.58. 2
~ ~ ~
2.4 2.0 CD C) 1.
6'.v) 1 CD
- 8
.4 0 4.0 8.0 12.0 16.0 .: '.-::20.0 CROSS FLOM VELOCITIES ft/sec---
FIGURE 4.6;.5-.2:. ~FLUIDELASTIC STABILITY RATIO OF TUBES AS A
'FUNCTION OF CROSS FLOH VELOCITY, FIXED-FIXED NDARY CONDITIONS 4.5-66
FIXED-PINNED BOUNDARY CONDITIONS DAMPING RATIO'='0.01
'YMBOL CROSS SECTION '1ST MODE FRE . -'Z 0 CIRCULAR .40. 3 Q FLAT ', 31.6 0 10K OVALIZED -'40 t
0 KIDNEY '39.9
..2;4 u Z.O>
CO 1.6
'UNSTABLE STABLE
--8
~ 4
'-0 e 4.0 8.0 ~
12.0 16.0::,,: -: '20.0 CROSS FLOW VELOCITY - ft/sec I ~ ~
....@FIGURE 4..5. 6"3: 'FLUIDELASTIC STABILITY RATIO OF TlSES AS A.
FUNCTION OF CROSS FLOW VELOCITY,'IXED-PINNED
'OUNDARY CONDITIONS
FIXED-PINNED'BOlJNDARY CONDITIONS FOR FLAT CROSS SECTION SYMBOL DA!<PING RATIO ~ ~
0 0.910 cf 0. 025 ~
4
~- C. i ~
4
, ~
4 2.4 a) 2,0 1.6 C/)
CD cCQ lD UNSTABLE LI STABLE
) c ~
~ ~ g 4
- .
- ,.4 0 4.0 8.0 12.0,, 16.0, ',-.20 0 CROSS FLO< VELOCITY - ft/sec
~ 'I
.4.5-6-'4: WLUIDELASTIC STABILITY RATIO eONPARISON FOR DA!IPING
'."IGUANA RATIOS OF 0.01 AND 0.025 AS % F'UNCTION OF CROSS'LOH VELOCITY, FIXED-PINNED FLAT DISTRESSED TUBE
- 4. 5-'$8
0 I
, I a t
'YLINDRICAL CROSS SECTION TUBE
~ (
! I
~
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I SY%0L AXIAL FORCE FIXITY I-Q, l
0 ,0 F, F.
a
'F-P
&1000 1bs COMP F-P
~ ' a
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h
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-4.0 .'
---'8.0 - -'-12.0 :-'6.0'='"--.='+0.0
~ =
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' 'CROSS-FECAL-VELQCI.TY- ft/sec -':~ '-:-'-.
! ~
~ ~
- i. t FIGURE 4 ~ 5 ~ 6-5 r, .'PEAK RNS AMPLITODE INE TO TURBULENCE AS A ~ r
~ 4 4 I 'FUNCTION 8F:CROSS FLN VELdCITY I I
~ ~
I(aaaar llaaat
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5
. FLAT CROSS SECTION TUBE k I
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SYMBOL AXIAL FORCE FIXITY 'I
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CROSS.SLOW 5)ELOC ,'c/sec~II4'-," ,
rg JFIGURE 4 5 6~ . PEAg QF.
I ~J RMS AMPLITUDE DUE TO TURBULENCE CROSS a
FLOW .VELOCITY I, 5
-;,":.i-:
~ ~ L AS-4 fUNCTION
~ ~ J s s I
~ '
TORQUE LIFT
~. DRAG
~
FLUID VELOCITY U,;
f =SUID g ~ 2mf LIFT CL (pfU /2) HD exp (i~t)
FLARE 4.5.6I-7: .:Tear Nodel for Gall;ulat)ng DRAG = CD (pfU /2) HD exp (iDDt)
Vortex-Shedding Loads TORQUE CL (pfU /2) (HD /8) exp (iDGt)
~f Go 0 0.3 10. 0 20.0 VELOCITY ft/sec Lift vs. Ve1octty 100 '0 I
~ ~N
~ ~
ED e cr' QJ, 5 l eQ 0 0 0.0 '10.0 y".. )VELOCITY ft/sec Fi quency vs. Ye1ocfty FIGURE.4,5.6-S: OLternating Lift (Lbs) and Frequency (HZ) vs.'Luid VeLocity for a Tear 0 5 Inch Wide by 2.5 Inches Hi'gh
. L5-7'P
~ .0) 0.&
~.~)
~.~I
~ .a
~ .a ot.a VELOCITY ft/sec OR(UE vs. VELOCITY ZrCURZ 4.5.. 6-9 ALternating Torque (in"Lbs) vs. FLuid VeLocity For a Tear 0.5 Inch Wide by 2.5 Inches High
- 4. 5-73
10.0 20.0 VELOCITY ft/sec Drag vs. Ve1ocity 10.0 20. 0 VELOCITY ft/sec Frequency..vs. Ve1ocity ALternating Drag (Lhs) and Frequency (M7) vs. FLufd VeLocity for a Tear 0.5 Inch Wide Sy 2.5 Inches ILLgh 4..5 76
4.6 Model Test 4.6".1 Test Objectives A partial full scale flow model of the downcomer and tube bundle entrance region of the Ginna steam generator Number 4 wedge region was constructed and used for testing at the Westinghouse Engineering Test Facility in Tampa, Florida. The purpose of the test was to investigate the following elements associated with the postulated failure mechanism.
a) The nature and extent of foreign object mobility in the downcomer annulus.
b) The magnitude of foreign object impact loads on tubes.
c) The stability characteristics of degraded tubes near the tube bundle entrance region when subjected to various flow conditions.
d) The nature and extent of tube-to-tube interaction once a tube becomes severed near the top of the tube sheet.
4.6.2 Test. Apparatus The tube bundle at the flow inlet region of the Ginna steam generator was represented using forty-eight (48) tubes (.875 inch O.D.) extending from the tube sheet to the first tube support plate. The flow model is illustrated in Figures 4.6.2-1 through 4 .6.2-4 . The orientation of the test model with respect to the overall steam generator geometry is shown in Figures 4.6.2-5 and 4.6.2-6.
The tests were run at ambient pressure and temperature using the cold flow loop illustrated in Figure 4.6.2-7 . Water flow to the test model was controlled by a flow measuring venturi. The maximum downcomer velocity was 14 ft/sec during flow testing.
The flow field between the tube sheet and the first support plate con-sists of approximately 75 percent cross flow and 25 percent parallel flow for the region modeled. Therefore, the perforated rear boundary plate (Figure 4.6.2-3) contains three times the flow area as the first tube support plate.
A ENDEVCO Model M42A Biaxial Piezoelectric Accelerometer was placed inside selected tubes (R44 C58 and R45 C54) to sense tube motions resulting from impact forces and fluid flow. Refer to Figure 4.6.2-8 for accelerometer orientation details. The accelerometer was located in the tube midway between the simulated support plate and tube sheet and was connected, through a low noise cable and pre-amplifier to Unholtz-Dickie Model 1112H charge amplifiers/signal conditioners.
Output from these conditioners was recorded on a Tektronix Series 5000, 8 channel, storage oscilloscope with a Polaroid camera attachment.
Displayed data were in units of "g's" per division.
4.6-1
PCB Pieotronics impact force transducers were also mounted on perpen-dicular axes in the tube support plate for selected tubes in the test vessel (R44 C58 and R45 C54). Refer to Figures 4.6.2-9 and 4.6.2-10 for orientation details.
These sensors provided an output of 50 millivolts per pound force. Data acquisition was by the Tektronix oscilloscope with the Polaroid attach-ment. Displayed data were in units of pounds force per division.
During calibration testing, an instrumented impact hammer was used to generate tube accelerometer and force transducer time histories. The signatures derived from these known impact forces were compared with signals obtained as a result of the foreign object impacting with an instrumented tube during flow testing.
4.6.3 Test Procedure and Results A foreign object, with similar dimensions to the largest object removed from the Ginna B-Steam Generator, was positioned in the downcomer annu-lus of the test model. The object is illustrated in Figure 4.6.3-1.
The orientation and position of the object within the wrapper annulus region was varied to determine its relative stability and susceptibility to flow induced vibration. At each position, the object motion was observed through windows in the side of the test vessel and accelero-meter and force transducer time histories were obtained.
4.6.3.1 Object Nobility Object motion was random in nature and occurred for virtually all orientations and positions. The object demonstrated the ability to assume ability to 6.6.3.2 ~li various positons within the downcomer annulus as well as the porvide a relatively uniform cyclic loading on the tubes.
2 Foreign object impact forces were estimated by comparing accelerometer time histories recorded during flow testing with'hose obtained by striking the same tube with an instrumented impact hammer. Typical accelerometer and force transducer time histories recorded from foreign object impact during flow testing are shown in Figures 4.6.3-2 and 4.6.3-3. Acceleration and force envelopes are shown in Figure 4.6.3-4.
The maximum recorded tube acceleration was approximately 200 g and the maximum force transducer response was 30 lb It should be noted that the force transducer load will only be a small fraction of the impact load since the major portion of the impact load will be reacted at the tube
'heet end of the tube and only a minor part of the impact load will be reacted at the support plate end where the force transducer is located.
Typical accelerometer and force transducer signals recorded from impact with the instrumented metal hammer are given in Figures 4.6.3-5 and 4.6.3-6. Reasonable agreement with tube response signatures from foreign object impact during flow testing can be observed by comparing Figures 4.6.3-5 and 4.6.3-6 with 4.6.3-2 and 4.6.3-3. Data relating 4.6-2
hammer force, tube acceleration and force transducer response are plotted in Figures 4.6.3-7 and 4.6.3-8. Based on accelerometer calibration data, maximum foreign object impact forces measured during flow testing ranged between 120-180 lbs. Using the force transducer calibration data, maximum measured impact forces ranged between 200-350 lbs. The duration of the impact force was short, approximately one millisecond.
In order to investigate tube degradation from the foreign object during flow testing, an extended flow test was performed in which the object was allowed to remain in the test vessel for eight (8) hours, while the downcomer flow was maintained at 14 ft/sec. Visual examination of the tubes indicated degradation at several locations, with the most signi-ficant degradation occurring on the R45 C54 tube. A photograph of the resulting degradation is provided in Figure 4.6.3-9.
4.6.3.3 Stability Characteristics Tubes with locally degraded, structurally degraded, and/or severed cross-sections could experience flow induced vibrations leading to fatigue failure. In order to investigate the flow induced vibration characteristic of degraded tubes, the tube in R44 C58 was locally degraded by progressively machining away the cross-section at the location of maximum foreign object scars approximately four inches above the tube sheet face.
Tubes with undegraded and locally degraded cross-sections were stable with respect to flow induced vibration during all flow testing. Figure 4.6.3-10 illustrates the tube in R44 C58 with only 40 percent of the circumference remaining. Accelerometer response envelopes are illustrated in Figure 4.6.3-11 for 40 percent remaining circumference and 10 percent remaining circumference at 14 ft/sec downcomer flow without a foreign object being present. Tube accelerations were higher with only 10 percent of the circumference remaining, which is indicative of higher tube vibration amplitudes as expected. Accelerations became more erratic when the foreign object was introduced in the vicinity of the tube with only 10 percent of the circumference remaining, as shown in Figure 4.6.3-12.
4.6.3.4 Tube-to-Tube Interaction Severed tubes could degrade adjacent tubes through tube-to-tube impact and/or sliding which results in wear. To investigate the behavior of severed tubes at the- periphery of the tube bundle, the tube located in R44 C58 was severed at the location of maximum foreign object impact scars, approximately four inches above the tube sheet surface. The resulting accelerometer response curves for an accelerometer located within the severed tube, Figure 4.6.3-13, indicated intermittent impacts with adjacent tubes. The severed tube did tend to nestle between the R43 C57 and R43 C58 tubes. The observed amplitude of tube motion in this position was small, less than one-tenth of an inch.
4.6-3
4 ~ 6.4 Conc) us i ons The following overall conclusions were drawn from the flow model testing, and support the proposed mechanism:
a) Foreign object movement in the downcomer annulus was r andom in nature. The foreign object demonstrated the ability to assume a variety of positions in the annulus region.
b) Based on acceleromet'er readings, foreign object impact forces between 120-180 lbs are possible. Using force transducer data, impact forces between 200-350 lbs are possible.
c) Tubes with undegraded and locally degraded cross-section were stable with respect to flow induced vibration during all flow testing.
d) The tubing accel'erometer response envelope increased in magnitude as the tubing cross-section was progressively degraded by machining it away.
e) Tube acceleration increases for a degraded tube were more erratic when the foreign object was permitted to interact with the tube.
f) A tube sever ed near the tube sheet interacted intermittently with adja- cent tubes. The severed tube tended to nestle between neighboring tubes and in such a position the tube motion was small, less than one-tenth of an inch.
4.6-4
Viewing Mindow (Both Sides)
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+II II A
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'i Q Perforated Exit P1ate
.25 ~ 29 Figure 4.6.2-1, Plan View of Flow Test Model
- 4. 6-5
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4.00
~ r r 4
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l
~ Plate Support 4
I
.Perforate Exit Plat 66" 2.00
~ 8 y II 44 %8 32.38'ube I
I AO" l I
Figure 4.6..2-2 ~- Section View of Flow Test Model 4.6-6
~
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e Figure 4.6.2~3 Rear Elevation View of Flow Test Model Showing Perforated Exit Plate.'.G-7
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figure 4. 6 2-4 Photograph of Xbbe Bundle Used in Flow Test Model
- 4. 6-8
1 ..REGION OF TEST SIMULATION WRAPPER SUPPORT WITH WEDGES PATCH PLATE SHELL WRAPPER TU8E LANE Figure 4.6.2-5 4.6-9
FIGURE 4.6.2-6
~ Test Simulation Detail WRAPPER ROW 45 PATCH PLATE 34 SQUAR E le2 PlTCH TYP.
TUBES WRAPPER SUPPORT e (.are o.D.)
WlTH ROW 44 WEDGES ROW 43 ROW 42 ROW 41 SQLL ts 0 0 COL. 46 0 0 COL. 47 0 COL. 8 0 GUTLINE OF COL.49 TEST GEOMETRY 0 COL. 50 S COL. 5 l COL. 52 COL. 53 COL. 4 PLUG WELD S C L.55 .887 DlA.
ClRCULATlON COL.56 HOLES:
'8~o u COL 7 0
0 0
0 COL 62 4.6-10
6 MAIN FW IHLET TUBE SHEET GINNA TANGENTIAL 'FLOW TEST MODEL 4 2" RETURN TO COOLING TOwER SUPPLY FROM I CITY WATER tt T C@J NG 2750 GALLON I T*HK P
4 A&20 2 LOW VENTURI 2 I S6 GINNA FOREIGN OBJECT 4'P 2 IS PSIG SRV PI s 40 HP PRIMAHY PUMP
@V5 IO P4 VEST I SCHEMATIC 0-ISO V3 ORAwH
~ BY VENTURI I APP~~ REV i V1 0 250 Figure -4.6. 2-7 Cold Plow Loop
aAuiAi (<3 Hl-AXI A L p,cC E LERQMETER TAhlcxE t4'T lAD (T)
FLOW Figure '4. 6. X-8 Biaxial Accelerometer Orientation
Y-Force
- Transducer '/ t/
'-Force Transducer A
I I Flow picture 4.6.2~9 Force Transducer Orientation and Installation Assembly
- 4.6-13
0 Fjggze 4i 6<2 10 Force Transducers Mounted in Tube Support Plate 4.6-14
Thickness 1/2 inch Weight 3.51 lbs U
, Figure 4. 6. 3-1 Foreign Object (Full Scale)
- 4. 6-15
R Acceleration (100 g/div)
T Acceleration 50 MSec/Div R Acceleration (100 g/div)
T Acceleration 5 MSec/Div Figure 4.6.3-2 Tube (R45 C54)'ccelerometer response time histories for foreign object impact. Downcomer flow ~ 14 ft/sec 4.6-16
X Force (10 lb/div)
Y Force 50 MSec/Div X Force (10 lb/div)
Y Force 2 MSec/Div
~ g 4~ 6~3 3 Tube (R45 C54) force transducer time histories for foreign object imPact ~ D nc. ower omflow '14 ft/eec I
- 4. 6-17 '.'i
R Acceleration (100 g/div)
T Acceleration 2 MSec/Div X Force
'(10 lb/div)
Y Force 2 MSec/Div Figure 4 ~ ~ ~ 3-4 Tube (R45.C54) response envelopes for foreign object impact.
14 ft/sec downcomer flow.
- 4. 6-18,
Impact Force (20G lb/div)
R Acceleration (100 g/div)
T Acceleration 50 MSec/Div Impact Force (100 lb/div)
~e I
R Acceleration 4
p >)" (100 g/div)
T Acceleration 5 MSec/Div Figure 4. < 3-5 Tube (R45 C54) accelerometer response time histories for instrumented .impact hammer.
- 4. 6-19
Impact Force (200 lb/div)
X Force (10 lb/div)
Y Force 50 MSec/Div Impact Force (20'b/div)
X Force (10 lb/div)
Y Force 2 MSec/Div Figure 4 6 3"6-
~ ~ Tube (R45 C54) force transducer response for instrumented'.impact hammer
- 4. 6-20
400 300 A
0 I
IU tg 200 y e 8 4 M O~
o 100 r r
~ o Tube I
(R45 C54) g 0
100 200 300 400 Radial Acceleration, g Figure 4 4 3-2'ccelerometer Calibration Data.
4.6-21
400
~ ~
e 300 0
0 4J 0
g 200 Pl W yO ~
0 100 Tube (R45 054)
~t I
0 0
0 10 20 ,30 40 Resultant Transducer Force (lbs)
Figure 4..6.3-8 Force Transducer Calibration Data.
4.6-22
, pjgure 4.6.3 Tube degradation after 8 hours9.259259e-5 days <br />0.00222 hours <br />1.322751e-5 weeks <br />3.044e-6 months <br /> of flow testing with foreign object in downcomer annulus. Maximum degradation occurred at R45 C54 tube.
- 4. 6-23
Front View Side View
.~~re '4.6.3-10 Tube(R44 C58) with machined opening. 40% remaining circumference.
4.6-24
R Acceleration (1 g/div)
~ ~
T Acceleration 2 MSec/Div 40% Remaining Circumference R Acceleration (1 g/div)
S.qà rJ 'pl P'
s T Acceleration 3~. w ~
2 MSec/Div 10% Remaining Circumf erence
~gyre 4 6, 3-$ 1 Tube (R44 C58) accelerometer
~
response envelopes for 14 ft/sec downcomer flow with tube degration. No foreign object.
.4. 6-25
I ~ I I
s
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'I i )I (1 g/div)
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T Acceleration 2 MSec/Div Envelope R Acceleration (1 g/div)
T Acceleration 50 MSec/Div Single Sweep 1
Figure 4i6/3-12 Tube (R44. C58) accelerometer response for 14 ft sec downcomer flow with tube degradaticn(10% remaining cir-cumference) . Foreign ob)ect is present.'.
R Acceleration (1 g/div) l I T Acceleration II g f p @It 2 MSec/Div Envelope Cc R Acceleration (1 g/div) r I T Acceleration agore 50 MSec/Div Single Sweep t for 14 4,6.3-13 ft/sec Severed tube (R44 C58) accelerometer downcomer Slow. No foreign object.
response 4 5-27
I 4.7 }.aboratory Tests 4.7.1 ~Ci1 This test was designed to verify that repreated lateral load'impacts on a 0.875 inch outside diameter (0.050 inch wall) steam generator tube subject to a 1000 psig external pressure could ultimately result in collapse. The lateral loads are applied by an impact rod driven by a hydr aul ic cylinder.
4.7.1.1 ~Scimens The specimens are Inconel 600 steam generator tube with 0.875 inch O.D.
and 0.050 inch nominal wall. The test lengths are either 19.2 inches or 52.0 inches. Some specimens were machined to provide a 0.030-inch wall over the 4.00-inch test section being impacted.
4.7.1.2 Test. Equi pment The test fixture schematic and photographs of the test setup are shown in Figures 4.7.1-1 through 4.7.1-3. The photograph in Figure 4.7.1-3 shows the instrumentation details.
The tube specimen is mounted inside a 2.00-inch diameter tube with associated end caps, tees, seals, etc. A hydraulic cylinder applies the impact force through a rod and a hydraulic controller and function generator control the servovalve which controls the fluid supplied to the hydraulic cylinder.
The force, deflection, and acceleration are measured by the load cell, LVDT and accelerometer respectively. The LVDT and accelerometer can be seen in Figure 4.7.1-2. The transducer signals are monitored and recorded on the Nicolet storage scope.
The cyclic frequency and deflection are set to give the specified test load. At high frequencies, the hydraulic system cannot react fast enough to provide a large load. Therefore the frequency and/or deflection must be adjusted appropriately.
4.7.1.3 Test. Procedure The following procedure is used to perform the loading tests on the steam generator tubes. Prior to each test the specific test parameter s are determined and are used where the test procedure calls for "as specified" test requirement..
a) Cut tube to the "as specified" length.
4.7-1
0 b) Measure and record I.D., 0.0., and ovality of the tube in the test section. Mark the end of the tube to indicate the smallest diameter if there is ovality.
c) Install the tube in the test fixture ensuring the smallest diameter is in line with the impact rod.
d) Install end cap seals and other test fixture hardware.
e) Turn set screws to provide a hand tight contact with the tube.
f) Align the impact rod to meet the test requirements.
q) Apply 1000 psig water or gas (as specified) pressure. Check for leaks and eliminate if they exist.
h) Apply a 1000 lb axial load if specified.
Apply the lateral load, as specified, at various locations on the tube surface over an area centered four inches from the simulated tube sheet support, and app oximately two inches long in the axial direction and roughly forty degrees in the circumferential direction.
4.7.1.4 Test Results.and Conclusions The testing program is cur rently in progress and the results and conclusions will be reported in an addendum to this report.
4.7.2 ~Fi T The purpose of this test is to determine the fatigue characteristics of tubes structurally degraded such as seen in the Ginna steam generator.
Nominal 0.875 inch O.D. x 0.50 inch thick tubes will be mechanically degraded locally near one end and set in a tube sheet simulation (collar) at that end and a tube support plate simulation at the other.
The end conditions of the specimens will approach approximately a "fixed-pinned" situation. The structurally degraded section will be 2 inches long with its center 4 inches from the tube sheet simulated end.
Three damaged configurations will be considered: flattened (to simulate a full collapse), kidney-shaped (to simulate a partial collapse) and impacted (to simulate degradation as a result of interaction with a foreign object). The flattened shape is achieved by clamping the 2-inch section of the tube in a vise, and squeezing the tube. The kidney-shaped section is one in which half of the tube circumference is made to nestle into the other half. This is done by means of a specially-designed set of dies. The impacted cases will be one in which the tube will be degraded in a region centered 4 inches from the tube sheet to simulate impacting by a foreign object.
4.7-2
4.7.2.2 Test- Equi pment Figures 4.7.2-1 through 4.7.2-4 show the test setup being used.
The vibration exciter is attached to the tube with a small hose clamp approximately 10 inches from the pinned end of the tube. Oeflections are monitored with a linear variable differential transformer. The deflection is set by adjusting a micrometer to the specified (single amplitude) deflection. The tube is then vibrated until it just touches the micrometer . The LVDT output is then observed and maintained at that amplitude for the duration of the test.
An instrumentation block diagram is shown in Figure 4.7.2-5 .
4.7.2.3 Test Procedure Find fundamental bending resonance of tube.
b) Map mode shape of resonance. While monitoring response at one location on the tube to make sure the response level remains constant measure the response at approximately six increments along the tube. Use two miniature accelerometers for this purpose.
c). Fatigue, Test Adjust the maximum response amplitude to that specified in the table below. Run at this condition for 200,000 cycles. Failure will be detected as a drop in frequency. If failure does not occur proceed to the next step. Repeat until failure occurs.
~Am 1 itude ~Ste ~Am 1itude 1 .050 DA* 6 .200 DA 2 .070 DA 7 .250 DA 3 .090 DA 8 .300 OA 4
5,
.120 DA
.160 DA '09 .400
.500 DA DA d). Install a new specimen and run at next amplitude (or slightly lower amplitude) until failure.
Test 2 additional samples at 2 different amplitudes to attempt to get 3 well spaced failures.
4.7.2.4 Test Results Two tests were performed using the above p ocedures and are discussed below.
Dynamic or pea -to-peak amplitudes in inches.
4.7-3
A fixed/pinned beam with a 2-inch flat degradation was tested with no axial load. The natrual frequency was 39.29 Hz. The fatigue testing was performed with 240,000 cycles for Step 1; 283,000 cycles for step 9 and 361,000 cycles for step 10 or a total of 2.87 x 106 cycles. There was no failure.
A fixed/fixed beam with a 2-inch flat deformation was also tested with a 1000 lb axial load. The natural frequency was 64.72 Hz. The fatigue test was run with 202,000 cycles for Step 1 and 2.08 x 106 cycles for step 8. The amplitudes in the other steps were not used. No failure of the specimen was observed.
4.7.2.5 Additional Tests The following identifies additional planned testing.
a) Flat collapse: fixed-fixed, 1000 lb axial load and .005 inch deep by 1/4 inch long notches at 2 locations near the collapsed section.
The test will be run for a minimum of 106 cycles at .300" pA.
Kidney shaped collapse (1 inch long)
- a. Fixed-fixed: 0 lb axial load and 1000 lb axial load
- b. Fixed-pinned: 0 lb axial load
- 1) Run strain gage response survey.
- 2) Run fatigue tests if strain measurements warrant it.
c) Kidney shaped collapse 2 inch long same as 2. above d) Impact fatigue Vibrate tube such that it will impact against a hard edge measuring load at impact.
Run fatigue tests in one or more of the above three configurations as appropr iate.
4.7.2.6 Test. Results and Conclusions Testing is still in progress and specific test results and conclusions will be reported later in an addendum to this report.
ynamsc or pea -to-peak amplitude in inches.
4.7-4
0 AIR l5LEED NUT (TYP. SOTH ENDS )
CAP SCREW END CAP SPECIMEN R,O DIA. TUSE "O" RIhlCiS (TYP. BO'TN ENDS)
SWASELOK R.O DIA.TEE IOOO PSI WATER OR QA.EI DY'hlAMIC SEAL I M PAC'T AO'O S TANDOFFS LOAD CELL HYDRAULIC CYLINDER, I
I I
(
l4YDRAULIC CYLINDER MOUNTING FIX'TURE.
FIGURE 4.7..1-1:Test Fixture for Collapse'0f;Steam.Genera%of'ubes
~ g
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FIGURE 4. 71 2:PHOTOGRAPH SHOMING THE TEST SET-UP FOR TUBE COLLAPSE TEST 4~7q6
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FIGURE 4.7,f1. 3 PHOTOGRAPH SHONING INSTRUMENTATION DETAILS FOR THE TUBE COLLAPSE TEST
~~
U pz(zap 4.7.2-l. 'VIEW OF FATIGUE TEST SETUP o~ +
g
.kk"
~ %T
~~a 4.7.2-2 ANOTHER VIEW OF FATIGUE TEST SETUP 4.7.-8'
Q. e tP,,
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FIGURE .'4.7.2-3 CLOSE VIED OF FATIGUE TEST SETUP
.JC<'IGURE 4.7-2-4 ANOTHER CLOSE UP OF FATIGUE TEST SETUP
/.
- 4. 7-9 I
Accel erome ter Endevco Model 2220C Ampl i fier Kistler Accelerometer Model 504 Endevco Model 2220C Ampli fier Kistler LYDT Model 504
+150 HU RANGE Amplifier Fil ter CEC Ithaco Model 1-127 Model 4302 10 lb Vibration Exciter Fi1 ter Ithaco Model 4302 Power Amplitude Fi1 ter Oscilloscope Ithaco Phillips Model 4302 Model 3234 Vari able H.P. 5302 Os ci 1 1 ator Frequency Counter FIGURE4.7.2 5 INSTRUMENTATION BLOCK DIAGRAM
- 4. 7-10
4.8 Conclusions This section summarizes the major results of the various testing and analytical programs described in Section 4 of the report and based on these results, provides conclusions regarding specific elements of the overall failure mechanism which led to a rupture of tube Row 42 Column 55 of the B-Steam Generator.
The approach that will be taken in presenting the conclusions resulting from the steam generator tube failure analysis program will be to assimilate available information on the following specific elements of the overall failure mechanism:
Collapse Fatigue Wear Burst Pertinent information associated with collapse is summarized in Table 4.8-1. This information indicates that collapse of a plugged tube can occur as a result of pressure, axial, and lateral impact loads. It is concluded that plastic deformation due to normal operating loads, axial restraint loads, and lateral impact loads resulted in sufficient tube ovalization to cause collapse.
Pertinent information associated with fatigue is summarized in Table 4.8-2. This information indi cates that severing of a collapsed or degraded tube with a notch on stress riser can occur as a result of a continual lateral impact load.
interact with neighboring tubes Once a tube has been severed it can and through wear sufficiently degrade these tubes such that they could also internally pressurized.
fail in fatigue or by burst if Pertinent. information associated with wear is summarized in Table 4.8-3. This information indicates that wear played a significant role in propagation of the initial peripheral tube degradation toward the center of the tube bundle. Wear primarily seemed to be associated with a severed or free tube rubbing against another tube.
Pertinent information associated with tube bursting is summarized in Table 4.8-4. This information indicates that sufficient wear to result in bursting of an active tube can occur in a relatively short period of time and that the actual failure is consistent with analytical predictions.
Based on the above, it is concluded that the failure mechanism shown in was responsible for the January 1982 Ginna burst tube Figure 4.8-1 incident. The basic ingredients for initiation of the mechanism are plugged peripheral tubes and a foreign object. The foreign object, it is felt, interacted in an impacting manner with plugged peripheral tubes on the hot leg side of the generator and this ultimately resulted in
.extensive tube degradation and severance of a tube near the top of the tube sheet enabling the tube to pivot about a point of fixity at the first support plate. Such a severed tube could then interact with 4.8-1
neighboring tubes causing extensive tube degradation via wear and could ultimately experience a fatigue type severence at the first support plate.
Extensive tube degradation of neighboring tubes would result in their total failure and enable them to interact with other tubes.
In summary, it is concluded that the failure mechanism shown in figure 4.8-1 resulted in the January 1982 Ginna burst tube incident.
4.8-2
TABLE 4.8-1 COLLAPSE The conditions for tube collapse could be established by an external source wearing against the tube and thinning the tube over a portion of its surface and/or impacting a tube with sufficient magnitude to cause plastic deformation and tube ovalization.
Laboratory examination of a collapsed tubing surface identified significant cold work that could have been caused by a foreign object impacting against the tube.
Analytical calculations and model testing indicate that a foreign object similar to that removed from the steam generator is capable of impacting the tube with sufficient magnitude to cause plastic deformation.
Accelerometer readings taken during model flow testing indicate that the foreign object could exert impact forces in the range of 120 to 180 lbs. Force transducer data also taken during model flow testing indicated maximum impact forces in the range of 200 to 350 lbs. Analytical calculations indicate that impact forces in excess of 100 lbs are obtainable.
Analytical calculations for a 7/8 inch O.D. tube subjected to an external pressure of 1000 psi and a concentrated lateral load indicated that plastic deformation leading to tube ovalization will occur with lateral loads in the range of 50 to 75 lbs. The actual load required will depend upon tube ovality, wall thickness, and yield strength. An axial load could reduce the required lateral load for incipient yielding by approximately 10 lb.
4.8-3
TABLE 4.8-2 FATIGUE Fatigue failure of a tube can be caused by a significant reduction in its load carrying capability as a result of wear and/or by continuous impact by a foreign object.
Analytical calculations indicate that a nominal plugged tube with or without a notch or stress riser will not fail in fatigue considering worst case operating thermal and mechanical loads.
Structurally degraded tubes will not fail due to fluid induced vibrations alone.
Fluid elastic instability would not be predicted analytic'ally for a mechanically or structurally degraded tube. However, for a structurally degraded tube, the margin to instability is minimal.
During model flow testing tubes with degraded and structurally degraded cross section were stable with respect to flow induced vibration.
A structurally degraded tube continuously impacted by a foreign object can fail due to high cycle fatigue. Calculated failure times range from a few hours to a few weeks depending on the magnitude and frequency of the impact load.
Laboratory scanning electron microscope fractographs revealed fatigue type striations at failed tubing surfaces.
Oscillating lift, dr ag and torque fluid loads are not large enough to cause tearing of tubing protrusions, consequently tube shredding must be related to another process, such as wear.
4.8-4
TABLE 4.8-3 WEAR Extensive wear was instrumental in the propagation of initial peripheral tube degradation toward the center of the tube bundle and seemed to be associated with a severed or free tube r ubbing against another tube.
Metallographic examination of removed tube sections identified wear zones as having cold work surface layers indicating the presence of an external source such as a loose tube. Tube wear patterns were primari ly circumferential in nature extending over a fairly large length in the axial direction and were compatible with one-tube rubbing against another.
During model flow testing, a tube severed near the top of the tube sheet was observed to interact intermittently with adjacent tubes.
Analytical calculations indicate that sufficient degradation of a tube, due to wear of one Inconel tube on another can occur such that bursting of an active tube could result consistent with the actual Ginna data.
Eddy current indications in the B-Steam Generator hot leg periphery appear to be primarily of an I.D. type prior to February, 1979 .
Starting in February 1979, the pattern of O.D. indications identified by eddy current inspections is consistent with the development of 0.0. degradation by a foreign object in the most peripheral tubes followed by the propagation of degradation to neighboring tubes by a wear process, which could be caused by a neighboring tube severed near the top of the tube sheet and free to rotate about a point of fixity at the first support .plate or a piece of free tubing which had been severed at the tube sheet and first support plate.
4.8-5
TABLE 4.8-4 BURST Tube bursting occurred as a result of wear and is consistent with analytical predictions.
Calculations indicated that the minimum tube wall required to preclude bursting of a tube under normal operating conditions is 6.6 mils assuming a material ultimate strength of approximately 90 ksi.
This compares well with the 8 mils remaining wall measured in the laboratory for the burst tube.
Laboratory examinations indicated that the failure of tube R42 C55 was a purely ductile failure in the region of the tube wall that had been worn to a thickness of approximately .008 inches over a length of approximately 4.0 inches.
Analytical calculations indicated that sufficient degradation of a tube due to wear of one Inconel tube on another can occur such that bursting of an active tube could result consistent with the actual Ginna data.
4.8-6
Mechanical Load Active Tube Tube Plugged Collapse Vibration Shredding Sever Wear Wear Plugged Tube Active Tube Plugging Burst Figure 4. 8-1 Postulated Failure Mechanistn Sequence 4.8-7
5.0 STEAM GENERATOR REPAIR PROGRAM 5.1 Access Holes To provide access for detailed inspection and tube removal, two, 3 inch diameter access holes were drilled in the stub barrel on the secondary side of the B-Steam Generator. These holes were centered approximately 6 inches above the tube sheet on the hot leg side at the Number 4 and 6 wedge areas. Fig. 5.1 is a cross sectional view of the steam generator showing the vertical position of the access holes relative to the area of repairs. Fig. 5.2 and Fig. 5.3 are detailed plan views of the Number 4 and 6 wedge areas showing the location of the access holes relative to the tubes removed.
5.1.2 Installation The access holes were mechanically machined starting with a small pilot hole and enlarged with successively larger drills until the final, finished hole diameter was achieved. The holes were sized to preclude the necessity for weld reinforcement. Following repairs, the holes are sealed with a cover plate and gasket. The cover plate is attached to the steam generator by bolts threaded into holes drilled and tapped into the stub barrel. Fig. 5.4 shows the details of the cover plate assembly.
5.1.3 DesicCn The access holes are designed in accordance with the requirements of the ASME Code. A description of the structural analysis for the holes is contained in Section 6.2.1 of this report.
5.1-1
5.2 Tube Removal
- 5. 2. 1 ~Summa In order to determine the extent of repairs required to the B-Steam Generator, the defects in periphery plugged tubes were categorized as follows:
a) structurally degraded tubes that are collapsed, severed, or have visible through wall defects; the size of the defect is large enough to cause a significant reduction in section modulus.
b) video O.D. indication - tubes that were observed by video inspection to have minor dings, wear marks, or similar small defects on the outside; the defects do not involve any significant loss of volume or reduction in section modulus.
c) eddy current signal - tubes that were observed by video inspection to have no defects on the outside; the tubes have no defect beyond that for which they were originally plugged.
d) preventatively plugged - 3 tubes surrounding R42 C55 preventatively plugged prior to performing the secondary video inspections; these tubes showed no defect either by eddy current or video inspection.
e) pulled tube - one tube (R45 C52) was pulled for metallurgical examination in 1978.
All of the defects categorized as structurally degraded or video O.D. indication were in the section of tube between the tube sheet and first support plate in the hot leg side.
5.2.2 Cate orization of Tubes Figure 5.5 is a listing of all the plugged tubes by category and area on the B-Steam Generator periphery. There are a total of 56 tubes including R45 C52. Of these, 24 are categorized as structurally degraded (including R42 C55), 11 as having video O.D. indications, 17 as being plugged for eddy current indications, and 3 preventatively plugged. The number in each category by area is shown in the figure.
5.2.3 Tube Removal All tube sections categorized as structurally degraded have been removed from the steam generator. This included 6 from the Number 6 wedge area, and 19 from the Number 4 wedge area. The basic removal procedure (except for the pulled tube) involved cutting the. lower 8 to 12 inches above the tube sheet of each tube using an electrode discharge machining (EDM) procedure. The 5.2-1
remaining 38 to 44 inch section below the first support plate was
'cut using a mechanical cutter. The cutting operations were performed entirely through the access holes at the Number 4 and 6 wedge areas. The tubing sections were manually removed from the steam generator either through the access holes or the existing 6 inch handhole at the end of the tube lane.
5.2.4 Severed Tube Removal As indicated in Section 3.4, 2 tubes (R45 C54 and R44 C56) were found already severed at the tube sheet and first support plate.
These were simply cut into smaller sections and manually removed through the access hole at the Number 4 wedge area. The severance at the upper end, in one case, was just below the support plate and, in the other case, just into it. An analysis of this con-figuration is provided in Section 6.2.3.3 of this report.
5.2.5 Tube Pull One tube (R45 C47) with an eddy current indication approximately 24 inches above the tube sheet was removed for metallurgical examination. This tube was pulled from the secondary side of the steam generator. This operation involved cutting out a section of the tube in the U-bend region above the upper support plate; mechanically cutting the tube just above the tube sheet; hydraulically pulling the tube from above; and cutting was removed. The final step was it installation into sections as of a restraining it device on the cut tube end at the first anti-vibration bar.
Access to the U-bend region for this tube removal operation was provided by a manhole cut in the wrapper. This hole was sealed following completion of the pulling operation.
5.2.6 U-Bend Restraint During a similar pull of another tube (R45 C52) in 1978, U-bend and sections of 2 .adjacent tubes (R45 C53 and R45 C51) were also removed. Since the lower section of R45 C53 was removed during the present repairs, it was restrained at the upper end by expansion on either side of the support plate. This restraint is designed to preclude any possibility of this tube moving during operation.
5.2.7 Remainin Tubes than the removal of R45 C47 for metallurgical examination,
'ther no sections of tubes categorized in Figure 5.5 as having only video O.D. indications or eddy current signals were removed from the steam generator. A structural evaluation of tubes with minor O.D. defects is provided in Section 6.2.3 of this report. The tubes with eddy current indications were plugged in accordance with the requirements of the plant Technical Specifications.
- 5. 2-2
5.3 Loose Parts Removal As .indicated in Section 3.5, several loose parts and pieces of tubing were found in the steam generators. The few, small loose parts in the A-Steam Generator were simply removed manually or with the aid of remotely operated mechanical grippers. A more extensive program of inspection and removal was necessary in the B-Steam Generator. Where loose parts or pieces of tubing were accessible from the tube lane handholes, or periphery access holes, they have been removed manually. Where size allowed, they were removed with remote mechanical devices using grippers or magnets. Pieces too small to remove by these means have been removed by vacuuming or water lancing. These operations were designed to remove all loose parts and pieces of tubing, regardless of size, from the steam generators.
5.3. 1
5.4 Mechanical Plu Removal Since the damaged section of R42 C55 was removed from the 8-Steam Generator, the 3 adjacent good tubes (R41 C55, R42 C54, and R42 C56) have been restored to service. The 6 mechanical plugs installed in the tube sheet ends of these tubes have been removed.
The removal was accomplished by hydraulically stretching and pulling the plug from each tube end. Subsequent eddy current examination will be performed to assure that there are no unacceptable indications in any of these tubes.
- 5. 4-1
5.5 Material Control Several material control procedures were used during repair and
~
modifications of the steam generators. A Material Control Log
~ ~ ~
was used to verify that all material and tools entering the steam generator were either removed or installed. Openings around work areas on the secondary side were sealed to prevent material or tools from entering other areas of the steam generators. Lanyards were attached to parts being cut out to assure their retrieval.
In addition, parts removed during modifications were reconstructed outside the steam generator. Finally, Quality Control surveillance, hold points, and cleanliness inspections were used to verify that no tools or loose parts had been left in the steam generators.
5.5-1
5.6 Post Re air Ins ections ests Following completion of the repairs in the B-Steam Generator, a series of inspections and tests will be performed to assure that it is ready for return to service. Sections of tubes adjacent to areas involved in the repair operations will be eddy current examined to assure that no unaccepable defects are present. A final series of video inspections will be performed to assure that no loose parts or tubing fragments remain and to verify that no unacceptable O.D. defects were caused by the repairs. A secondary side hydrostatic test will be performed to verify integrity of the access hole covers. A primary side hydrostatic test will be performed to assure that no measurable primary to secondary leakage is present.
- 5. 6-3.
5.7 Radiation E osure 5;7.1 P~lannin Ginna Station's ALARA program was applied to all of the steam generator inspection, repair, and modification activities. All inspection, repair, and modification activities were reviewed technically and radiologically before implementation. Procedures and equipment were tested, and personnel trained, in full scale mockups prior to use in the steam generators. The secondary video inspection system was developed at Ginna Station using Rochester Gas and Electric's steam generator mockup. Special mockups were built to check out the equipment for machining the 3 inch access holes.
5.7.2 Tube Removal Due to the potential exposure associated with removal of the tube sections, a significant effort went into exploring various alter-natives. Several different types of tube removal methods and equipment were evaluated. The process finally selected provided a technically acceptable means of tube removal with the lowest radiation exposure. The process incorporated equipment from several sources. A Westinghouse-developed electrode discharge machining procedure (EDN) procedure was used for cutting the lower section of each tube. Equipment furnished by Babcock S Wilcox, NUS, and Alliance Tool and Die was used for the mechanical cutting and removal of the remaining tube sections. Development of this I.b. mechanical cutting and removal system eliminated the need for additional access holes or U-bend removal. This saved several hundred man-rem.
5.7.3 ~Ex osure The radiation exposure associated with inspection and repair of the B-Steam Generator will be approximately 310 man-rem. This includes approximately 200 man-rem for the initial inspections and Phase I repairs; 80 man-rem for tube removal; and 30 man-rem for the post repair inspections and tests. These numbers include the exposure associated with the inspections and repairs resulting from the tube rupture. They do not include the exposure received during the normal inspections and modifications performed, as previously planned, as part of the refueling outage. Table 5.1 summarizes the exposure experienced during this outage.
5.7-1
TABLE 5.1 1982 B-STEAM GENERATOR RECOVERY EXPOSURE Activit Man-rem Initial investigation of B-Steam Generator. 53.668 (a)
This involved opening both primary and secon-dary sides of the generator, ECT of hot and cold legs, plugging 21 tubes, sludge lancing and closing the primary side.
Fiber optic and video-camera examination of 32.209 (a) the secondary side of the generator.
Procedure EM-306. This involved second sludge 37.628 (a) lancing, opening and hydro-lasing the steam spaces, cutting the wrapper, third sludge lanc-ing, and setting up the platform and water shielding on the handhole area.
Drilling the first 3 inch port, in the Number 4 14.533 (a) wedge area.
Cut and remove tube samples through the 3 inch 8.841 (a) port (EDM cuts);
Shield the U-tube bundle, cut and remove 1 tube 21.645 (a) from the top.
Drilling the second 3 inch port, in the Number 6 13.097 (a) wedge area.
Fiber optics and video-camera evaluation of 2.923 (a) generator prior to closing.
Close out for generator removing lead from the fill.
U-tube Involves bundle, modify-.
11.340 (a) ing the seal scaffold, putting the wrapper cut-out in place, and closing the secondary side.
Following partial completion of secondary .800 (a) modifications, set up for lower removal of damaged tubes between the tube sheet and the first support plate.
EDM and router cuts and removal of damaged 23.000 (a) tubes in the Number 4 and Number 6 wedge areas.
Machining of a radius on the inside of the 1.470 (a)
Number 4 and Number 6 wedge holes.
Vacuuming, fiber optics and camera inspec- 1.500 (a) tion of the B-Steam Generator tube sheet.
5.7-2
TABLE 5.1 (Continued)
Activit Man-rem
- 14) Mechanical plug removal set up. 3.B35 (a)
- 15) Mechanical plug removal. 15.650 (a)
- 16) Stabilization of U-Bends and unsupported tube. 12 (e)
- 17) Eddy current test periphery area. 12 (e)
- 18) Alignment check laser. 2 (e)
- 19) Reinstall wrapper patch. 15 (e)
- 20) Water lance. 7.5 (e)
- 21) Final video camera inspection. 6 (e)
- 22) Final closeout. Includes removal of scaf- 10 (e) folding from steam space, and closure of all channel head and secondary openings, dismant-ling platforms and removal of water shields and lead.
Total 306.7 (e)
(a) - actual (e) - estimated 5.7-3
SLjPPo<T PLKTE
%HELL TUME. NHKC.T QOLt IHITIAI 0 oRIcINaL DATE DRAWN CHECKED It CSI', CHC,
~ CII RE V ISION OY ~Y CHC NANO'It ItOCHESTER GAS 8c ELECTRIC CORP. maCE AOCHE5TCIC, HKW YCNRK >" <CC.Ma +Ol t NO.
ACC SS HOLE GINNO STATION B-STEAM GENERATOR N0.4 WEDGE AREA SHELL rreo ~ eO 0 0 0
0 41 40 COLUMN 62 61 60 59 58 57 58 55 54 55 52 51 50 48 48 4l 48 45 0 FLtNt HOLES " M~ BURST/ REMOVED/PLUGGED DEFECT/ PLUMED NO DEFECT/ PLUGGED
( i PULLED/PLUGGED ACTIVE iP FIG.5.2
1 0
NNA STATION
-STEAM GENERATOR
- 0. 6 SEDGE AREA SHELL ACCESS HOLE 0
0 0
ALCOVE E'LAQUE INITIAI P ORIG INAI DATE ORAWN CNKCKEO RCSP, CHC.
~ CR REVISION SY ~Y INC, MAHC'R.
ROCHESTER GAS tk ELECTRIC CORP. G ENKZATCZ SCALE AOCHKSTEtt, HEW YCÃtK ACCESS ~LE. CPM NO.
FIGURE 5 GINNA STATION B-STEAM GENERATOR CATEGORIZATION OF DEFECTS NO. 6 R40C70 NO. 4 NO. 2 CATEGORY- WEDGE AREA AREA WEDGE AREA WEDGE AREA
- 1. Structurally R8C92 R42C5 5M R43C59 R44C55M Degraded R10C91 R43C53 R43C60 R44C56 R11C91 R43C54M R43C61 R44C57 R12C91 R43C55M R44C52 R44C58 R14C90 R43C56M R44C53M R45C53 R15C90 R43C57 R44C54 R45C54 R43C58
- 2. Video R9C91 R38C71 R45C51 OD Indication R13C90 R38C72 R16C89 R39C68 R17C89 R39C69 R39C70
- 3. Eddy Current R15C89 R35C75 R45C46 R12C2 Signal R40C67 R45C47M R28C12 R40C68 R45C48 R30C15 R41C66 R45C49 R31C15 R45C50 R32C15 R32C16 R33C15
- 4. Preventatively R41C55 Plugged R42C54 R42C56 TOTALS 28 M Metallurgical Samples R45C52 pulled April 1978.
6.0 Technical. Basis for Re airs 6.1 Introduction This section provides details of an analytical evaluation to verify the acceptability (and compliance to the ASME Code requirements) of the Ginna 8-Steam Generator. The pertinent analyses and evaluations are described in the following sections.
6.0-1
6.2 Analyses The following analyses were performed in support of the repair program:
a) Access holes: Structural analyses of the secondary shell for the two 3-inch diameter access ports were performed to verify the Code acceptance of the shell subject to the applicable loading requirements in the Equipment Specifications. Based on these the access ports are structurally acceptable. 'nalyses, b) Thermal/Hydraulic evaluation: Thermal/hydraulic analyses were performed to determine the type of flow redistribution due to the removal of structurally degraded tube spans on the hot leg peri phery and to examine the effect of the flow redistribution on the fluid induced vibr ation characteristics of the remaining tubes. Based on detailed analyses, it is shown that the maximum tube gap velocity is increased by approximately 12 percent; however, the modification is structurally acceptable since significant margin exists for both the fluid-elastic stability ratio and vibrational stresses due to cross-flow turbulence.
c) Structural evaluation: An evaluation was performed to address the structural acceptability of plugged tubes with visual surface irregularities such as small scars or stress risers. The evaluation considers both the fatigue margin under operating transients and collapse integrity of such a tube. Additionally, the geometric stability of, tube(s) severed just below the first tube support plate is also examined. Results of the evaluations indicate that plugged tubes with minor surface irregularities and tubes severed just below the first support plate are structurally acceptable and can be left in the tube bundle.
Details of the above three analyses are given separately in the following subsections.
6.2.1 Access Ports A structural analysis of the 3-inch diameter access ports was performed. The ports are located on the secondary side of the steam
= generator just above the tube sheet secondary surface. The location of the two ports is shown in Figure 6.2.1-1.
The geometry of the access port was modeled using the WECAN computer code, Reference 6.2-1. Three-dimensional isoparametric elements were "
used for the shell and cover and beam elements were used for the bolts.
Plots of the finite element model are shown in Figure 6.2.1-2.
The loadings and loading conditions were obtained from the Ginna Equipment Specification. The specific loadings applied to the finite element model include internal pressure, thermal transients, tube sheet interaction with the stub barrel, and bolt preload. The stresses from
.these loads were combined per the E-Spec requirements and compared to the ASHE Code Section III allowable values.
6.2-1
The results of the analysis are summarized in Table 6.2.1-1. As shown in the table, all of the ASME Code requirements are met. The design of the 3-inch diameter access ports is therefore structurally acceptable.
6.2.2 Thermal-Hydraulic Evaluation 6.2.2.1 ~Pur ese The removal of tubes from the periphery of the hot leg tube bundle between the tube sheet and the first support plate will lead to flow redistribution. The flow velocities resulting from this flow redistribution must be evaluated to determine if flow induced loads on the tubes around the tube removed region are significantly affected.
The CHARM computer program (Reference 6.2-2) was used to perform a thermal-hydraulic analysis in the region of the tube bundle between the tube sheet and the first support plate. Three cases were considered:
a) nominal, which will be the same as previously presented in Section 4.5.2, b) one block of tubes removed from the periphery of the bundle, and c) two blocks of tubes removed in the periphery of the bundle.
Case b is most representative of the proposed tube removal (= 80 tubes);
Case c may be considered a very conservative upper bound. The CHARM analysis was performed in the plane of symetry perpendicular to the tube lane which divides the hot and cold legs into equal halves. At nominal conditions, there is no flow across this plane.
Because the CHARM analysis was two-dimensional (axial and radial), a separate three-dimensional hydraulic analysis was performed of the tube sheet-to-first support plate region with the WECAN hydraulic conductance element (Reference 6.2-1). The purpose of this analysis was to determine the effect of tube removal on the three-dimensional aspects of the flow distribution. This analysis used a pressure-forced boundary condition.
6.2.2.2 Analysis The CHARM model used in the present analysis is basically the same as that described in Section 4.5.2 (see Figure 4.5.2-1). The only difference is that, to simulate tube removal, the flow resistances in the tubes-removed region were set equal to zero as well as setting the heat loads to zero. Thus, in columns two (and three for Case c) both the heat input and flow resistances were set equal to zero, whereas in columns 20 (and 19 for Case c), only the heat load was set equal to zero. (See Figure 4.5.2-1.)
The WECAN hydraulic conductance is basically the same as that discussed in Section 4.5.2. For those flow paths which connect the tube-removed region to the other parts of the model, the hydraulic conductances were altered to reflect the fact that the removal of tubes effectively decreases the flow resistance to zero over that part of the flow path.
As in the plugged tube case of Section 4.5.2, the deadwater region above the wrapper opening was simulated by setting the axial hydraulic conductances in this region to very small numbers.
6.2-2
6.2.2.3 Results Figures 6.2.2-1 through 6.2.2-4 show the CHARN lateral and axial velocity values and velocity vector plots for the one block of tubes removed and two blocks of tubes removed cases, respectively. As discussed in Section 4.5.2, the axial velocities shown in these figures are between tube axial velocities and need no correction. To convert the lateral or crossflow velocities to between-tube maximum velocities it is necessary to multiply the values shown on the figure by a factor of 2.08 to remove the porosity correction and account for the crossflow blockage presented by the tubes. For conservatism, an additional area ratio factor can be applied to account for the flow area convergence of the radially inward flow.
The 3-D WECAN analysis indicates behavior in the tube-removed region similar to that of the plugged tube region of Section'4.5.2. However, increased hydraulic conductances in the tube-removed region results in a slight increase in the radial flow velocities entering that region, unlike the plugged tube case, where the entering radial velocities decreased slightly below nominal values. Estimates of maximum crossflow velocities therefore included this effect as well as the effects discussed in Section 4.5.2. Otherwise the methodologies are the same.
Table 6.2.2-1 gives a summary of the maximum between-tube crossflow velocities for tubes. on the periphery of the tube-removed region as well as the face b crossflow velocities of several regions downstream (radially inward).
Figures 6.2.2-5 and 6.2.2-6 show the quality distributions predicted by CHARM for Case b and c. As in the plugged tube case, the main effect here is a shift of the vaporization boundary radially inward and away from the low velocity zone. Conditions just above the tube sheet are quite similar for the nominal, plugged tube and both tube-removed cases.
In summary, the removal of tubes from the peri phery of the tube bundle has the following effects:
The major 'effect of tube removal is the appearance of a reduced fluid velocity field and a low quality region between the wrapper opening and the first support plate in the region where the tubes are removed. This behavior is similar to the plugged tube case in Section 4.5.2 . This low velocity region tends to increase fluid crossflow velocities in the region near the wrapper entrance.
The highest between-tube crossflow velocity increases from 9.01 ft/sec in the nominal case to 10.12 ft/sec in the one block of tubes removed case and 10.95 ft/sec in the two blocks of tubes removed case. These are increases of 12.3 percent and 21.5 percent, respectively. Based on analyses presented in Section 4 .5, these increases have practically negligible effect on tube loadings and resonses.
6.2-3
6.2.3 Structur al.Evaluation The basis of steam generator repair has been to remove all structurally-degraded tubes. This includes tubes with significant reduction in stiffness and frequency due to the loss of cross sectional moment of inertia resulting from collapse, large structural discontinuities in the form of visual notches and cuts, and large holes. On the other hand, tubes with visual surface irregularities due to small ovality and distortion were not removed.
This section addresses the structural acceptability of the tube bundle followjng the repair effort. Specifically, the following two considerations are examined for a plugged tube with surface irregularities:
a) Fatigue margin under operating transients, and b) Collapse integrity Additionally, the geometric stability of a tube severed and/or cut just below the first tube support plate (TSP) is also verified. As far as surface damage on an active tube is concerned, it is to be noted that safe operation for an active tube is assured by eddy-current testing in accordance with the applicable tube plugging margin per the Technical Specification limits.
Consistent with the repair program objective, it is assumed that all structurally damaged tube spans between the tube sheet and the first TSP are removed as well as all foreign objects. For fatigue evaluation, therefore, only tubing with surface irregularities in the form of cross section distortions and/or surface irregularities need be considered.
As far as thermal mechanical loading on such a tube is concerned, the loads are the same as in the case of a nominally plugged tube described in Section 4.5. Effect of small distortion/ovality would be to increase the hoop bending stress due to the external pressure loading. For example, the maximum hoop stress for a nominally plugged 0.050-inch wall tube under 1000 psi pressure increases from 15.0 ksi at two percent ovality to 25.0 ksi at six percent ovality. This increase. in the hoop stress has a relatively insignificant impact on the usage factor.
Assuming a 30-year remaining plant life for a six percent oval tube with a design minimum wall of 0.045 inch, a significant fatigue margin exists based on the code calculations in Section 4.5.7 for a notched, plugged tube. Thus, from the view-point of fatigue due to operating plant transients, plugged tubes with surface irregularities are acceptable.
From the viewpoint of hydraulic loading, the effect of tube removal on the maximum tube gap velocity was considered in order to determine the acceptance of such tubing subject to the mechanisms of fluid-elastic
.stability, turbulence and vortex shedding. Oetai ls of thermal-hydraulic 6.2-4
analyses to determine the maximum gap velocities for the before and after tube removal cases were given in the previous section. Table 6.2.3-1 summarizes the results of these analyses. The one block of tube removed case corresponds closest to the post-repair steam generator tube bundle geometry. The expected maximum gap velocity is 10.12 ft/sec, or approximately 12 percent greater than the calculated velocity prior to the tube removal condit'ion.
Based on the results of flow-induced vibration analyses in Section 4.5.6, it is seen that the calculated velocity change has a rather
.insignificant impact on the fluid-elastic stability, turbulence and vortex shedding responses of a tube with given cross section and under a given boundary condition. The results also indicate that response of a tube with small ovality and/or distortions are essentially the same as those of a nominally round tube. The actual vortex shedding and cross-flow turbulence amplitude of a fixed-fixed tube span with various degrees of localized distortion and subjected to a 10.0 ft/sec cross-flow velcotiy over the 14 .0-inch wrapper opening are summarized in Table 6.2.3-2. Again, the comparison indicates that with the exception of the case of si gnificant tube distortion or collapse, the vibration amplitudes are r'elatively stable, that is, about the same as the nominal round tube.
6.2.3.2 Collapse Inte rity Inconel-600 tubing typical of PWR steam generators has been extensively tested to determine the effect of local degradation on the external collapse pressure strength of the tubing. Figures 6 .2.3-1 through 6.2.3-3 show the results of an NRC-sponsored test program, Reference 6.2-3. These tests. were performed on mill-annealed Inconel-600 straight length tubing with nominal 7/8 inch 0.0. x 0.050-inch wall placed in a 600 F autoclave, and pressurized externally by water simulating the secondary side steam generator fluid. Local degradations considered in this testing include: (1) elliptical wastage typical of local wear due to tube-to-tube or tube-to-a-foreign object contact, (2) uniform thinning typical of chemical wastage, and (3) EDM slots to simulate axial cracking.
The following is to be noted, based on these test results:
a) Expected collapse strength of a nominal tube is approximately 5000 ps .
1 b) C'ollapse strength is relatively unaffected by short (length less than or equal to the tube diameter) through-wall cracks.
~) For tube collapse corresponding to the external pressure of 1020 psi (maximum expected secondary side pressure for the Ginna steam generators) required tube wall degradation is approximately 80 percent for uniform thinning, and greater than approximately 90 percent for localized thinning.
Thus tubes with small surface scars and localized wear have significant margin to collapse. As far as the effect of local tube distortion/
6.2-5
ovality is concerned, it is to. be noted that 1) the collapse mode of tube failure results from plastic instability of the tube shell and thus, represents an instantaneous failure mode, 2) of all the design-basis loading conditions for the Ginna steam generator tubing, the maximum secondary side pressure of 1020 psi occurs during normal operation at the hot standby condition.* In other words, plant operation (at hot standby) in itself represents a proof collapse test.
Therefore, tubes with "local distortions have ovality below the threshold of plastic instability and consequent collapse. In the absence of any external mechanism, these tubes are expected to remain stable during subsequent oper ation.
6.2.3.3 Geometric Stability of Cut/Severed Tubes The degraded tube sections between the tube sheet and the first TSP are removed by cutting the tube spans near the top of the tube sheet and 2 to 4 inches below the first TSP. Additionally, a small number of tubes had their lower spans severed (due to fatigue) just below the first TSP . It is imperative that the geometric stability of the remaining partial tubes, as schematically shown in Figure 6 .2.3-4, be assured by verifying that the broken legs will be confined within the TSP.
The broken leg of the tube can move axially with respect to the confining TSP as a result of both the mechanical pressure loads in the tubing as well as thermal growth mismatches. The following summarizes the worst case loadings and relative motions.
o Worst Case Pressure Loading Subject to the 2560 psi maximum primary pressure differential during a postulated feedline break condition, the maximum broken leg motion, 6, is calculated assuming the broken tube is coupled rigidly to the tube bundle at the U-bend. Using the largest radius tube bend,
< = P AFL/AmE = 0.14 in where p = inside pressure differential Af = flow ar ea inside the tube Am = tube metal area L = length of largest tube E = Young's modulus The maximum calculated stretch of 0.14 inch is significantly less than the TSP thickness of 0.75 inch. Hence, the broken tube end will remain confined within the TSP.
- Secondary side hydro test, although at a pressur e somewhat higher than the hot standby pressure of 1020 psi is not critical due to the offsetting effect of higher yield strength at the lower temperature of secondary hydro.
6.2-6
o Thermal Growth Mi smatches, Thermal growth mismatches due to tube-to-tube and tube-to-shell interactions during various thermal transients can result in motion of the broken tube leg relative to the TSP.
For a total pull-out through a 0.75-inch thick TSP, a required thermal differential of T = 1900 F is calculated assuming a constant expansion coefficient at 600 F. Compared to the required T = 1900 F for the pull-out, the maximum expected aT during normal operating and postulated LOCA transients are less than 100 F and 400 F, respectively. Thus, a significant margin to pull-out exists due to thermal growth mismatches.
In addition to the above direct axial movements, the broken tube end can move axially also due to lateral tube deflection resulting from seismic and flow-induced vibrations. However, for the tube span between the first and second tube support plates, the vibration amplitudes are very small. Consequently, no significant axial movement of the broken tube will result.
6.2.3.4 Conclusions Based on the discussions above and the results of analyses presented in Section 4.5, the following conclusions are applicable to the post-repair structural integri ty of plugged tubes with slight surface irregularities in the form of small scars, local wear and distortion.
Fluid-elastic stability, vortex shedding and turbulence responses are practically unaffected by small distortions and surface irregularities.
Removal of tubes has no adverse impact on remaining tube stability due to fluid interactions. For surface degraded tubes, acceptable fatigue margin exists for subsequent operation. Plant operation being a proof collapse test, structurally stable tubes will remain stable during
'ubsequent operations. Tubes severed at the first TSP are geometrically stable and cannot pull out of the plate due to operating and faulted transients.
The safety and integrity requirements of active tubes are satisfied by existing Technical Specification limits for steam generator tubing.
6.2-7
e References for Section 6;2 6.2-1 WECAN Westinghouse Electric Computer ANalysis User's Manual, Second E3ition, March 1981.
6.2-2 A. C. Spencer, "Method of Characteristics for Solving Two-Dimensional Reactor Core Flows," Proceedings of the 1973 Conference on Mathematical Models and Computational Techniques for Analysis of Nuclear Systems, CONF-730414-PI(ANS), 1973, pp.
III-3 to III-22.
6.2-3 Vagins, M., et.al., ".Steam Generator Tube Integrity Program Phase I Report," NUREG/CR-0718, September 1978.
6.2-8
.TABLE 6.2.1-1 STRESS
SUMMARY
FOR 3 INCH DIAMETER ACCESS PORT Ratio- of Maximum. Stress to Al l owable. Stress Load. Condi ti on Bolt Cover Shel l Design 0.63(1) 0.29(3) (6)
Normal and Abnormal o.65(1) (1.0 1.O1(5)
O.98(2)
Test o.65(1) 0.29'(3) (6)
O.62(2)
Fatigue Usage Factor o.s5(4) 0.00 0.16 Bolt Replacement Interval 8 years Notes: (1) Average Service Stress (2) Maximum Service Stress (3) Primary Membrane Plus Bending (4) Fatigue usage factor based on specified replacement interval (5) Acceptable per Code. A simplified elastic - plastic analysis was invoked for the fatigue evaluation.
(6) Primary stress limits are satisfied by Code rules for opening not requiring reinforcement.
6.2-9
TABLE 6.2.2-1 BETWEEN-TUBE CROSSFLOW VELOCITIES IN AND NEAR THE TUBES RENOVEO REGION Crossflow Velocity Crossflow Velocity Crossflow Velocity (1 block of 2 blocks of Location (nominal .- tubes removed * - tubes removed) .
Perimeter Cell 8.2 ft/sec 9.13 ft/sec 9.13 ft/sec (Face A)
Perimeter Cel l 8.2 ft/sec 9.13 ft/sec 9.13 ft/sec (Face C)
Perimeter Cell 9.01 ft/sec 10.12 ft/sec 10.33 ft/sec (Face B)
One Cell in 8.54 ft/sec 9.48 ft/sec 10.95 ft/ sec from Perimeter (Face B)
Two Cells in 8.21 ft/sec 9.16 ft/sec 10.46 ft/sec from Perimeter (Face B)
- This case is most representative of the actual tube removal.
6.2-10
TABLE 6.2.3-1
SUMMARY
OF MAXIMUM TUBE GAP VELOCITIES WITH AND WITHOUT TUBE REMOVAL
-- ~
Case .-- Maximum-Gap.Velocity; ft/sec ,
Nominal 9.01 One block of tubes removed 10.12 (6.0 inches on periphery)
Two block of tubes removed 10.95 (12.0 inches on periphery) 6.2-11 rr
'TABLE 6.2.3-2
SUMMARY
OF VORTEX SHEDDING, AND TURBULENCE ANALYSES FIXED-FIXED BOUNDARIES CROSS-FLOW VELOCITY, 10.0 FT/SEC DAMPING RATIO, 0.01 Vibration. Amplitudes;.Mi ls.
Cross Section of Distorted Zone.- Vortex.Sheddin Turbulence Cyl inder (nominal ) 0.77 0.81 10 percent ovality 0.79 0.83 Kidney 0.79 0.83 Flat 2.13 1.53 Vibration amplitude due to vortex shedding and cross-flow turbulence are relatively unaffected by small distortions and surface irregularity.
- 6. 2-12
Figure 6 2.1-1: Location of Access Ports 6.2-13
Outside View Inside View Figure 6;4 .1-'2: Finite Element Model i
6.2-)4
NO RESISTANCE 6 HEAT LOADS
~ 15 .35 3.94 2.93 .86 1.93 1.56 1. 25
- 1. 36 1.83 : 1.31 .91 .74 3.94 3. 93 2.38 2.22 1.73 1.35 .97 .51 1.17 1.58 1. 12 .77 .61 .65 1.64 3.94 3. 94 2.96 2.25 1.75 1.38 1.02 .57 .17
.92 1.25 .91 .61 .48 .51 1.31 2.06 sec
- 3. 94 3.95 2.99 2.28 1.78 1.41 1.07 .62 -. 16
.68 .94 .68 .46 .35 .37 .99 1.58 3.94 3.96 3. 01 2. 31 1.81 1.44 1.12 .67 .16
.45 .62 .45 .30 .23 .24 .67 1.08 3.94 3.97 3.03 2.33 1.83 1.48 1.16 .72 ~ 15
.23 .31 .23 .15 ~ 12 .34 .56 3.94 3.98 3.04 2.34 1.85 1.50 1.21 .77 -.15 Lateral Velocity ft/sec +
Axial Velocity ft/sec +
X= 1 Fiaure 6.2.2-1 . RGE WITH FLOW RESISTANCE AND HEAT LOADS REMOVED IN 1 COLUMN OF CELLS (COLUMN 2) 6.2-15
NO RESISTANCE & HEAT LOADS
.38 .23 .69 .86
-.60 1.04 1.52 1.19
-.012 .77 1.15 1.19 1.09 .90
.85 1.26 1.58 1.19 .91 .88 3.94 3.04 2.76 2.24 1.81 1.46 1.12 .79 .32
.95 1.32 1.52 .82 .76 1.37 2.47 3.94 4.02 3.43 2.58 1.99 1.57 1 ~ 19 .86 .36
.81 1.14 1.29 .95 '. 69 .63 1.12 2. 09 fe sec 3.94 4.04 3.45 2.61 2.02 1.58 1.22 .89 .39
.64 .91 1.03 .76 .55 .49 .89 1.70 3.94 4.05 3.47 2.63 2.04 1.61 1.25 .94 .42
.47 .68 .77 .57 .41 .36 .66 1.29 3.94 4.05 3.49 2.65 2.06 1.63 1.28 .98 .46
.31 ,.45 .51 .38 .27 .24 .43 .88
- 3. 94 4 '6 3.50 2.66 2.08 1.65 1.31 1.02 .48
.16 .23 .26 .19 .13 .12 .22 0'.50
.45 3.94 4.06 3.51 2.67 2.09 1.67 1.33 1.06 Lateral Velocity, Axial 5 7ocity,,
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X= 1 2 Figure 6 2.2-2. RGE WITH FLOW RESISTANCES AND HEAT LOADS REMOVED IN 2 COLUMNS OF CELLS (COLUMNS 2 AND 3) 6.2-16
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REGION OF REGION OF TUBES REMOVED TUBES PLUGGED 1st TSP E OF
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FIGURE 6.2.2-6 ~
EQUALITY DISTRIBUTION WITH TUBES REMOVED IN COLUMNS 2 AND 3 AND PLUGGED IN COLUMNS 19 AND 20
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5 0
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0 lNDEFKCIED 0 0 WRAP ANGlE O 45o WRAP ANGlE l35o WRAP ANGlE 0
0 40 60 M MAXIMlNDEGRADATION, S wall FIGURE 6>2.3-1 Collapse Pressure of Tubing with Elliptical Wastage (from Reference 6.2-3)
- 6. 2-21
r9 o 0 UNDEKCTED 0 I/4" SLOT 0 1/?'SLOT
> 1-1/2" SLOT 0
0 N 40 80 MAXIMLlhDEGRADATION, % wall FIGURE 6.2.3-2 Co]lapse Pressure of Axially (EDM) Slotted Tubing (from Reference 6. 2-.3) 5 Pn 4 CC CL CL CP~
6 0 NDEFECTED 0 3/16" THINNED lENGTH 0 3/F'HINNED lENGTH
~ 3/4" THINNED LENGTH
+ 1-1/2" THINNED lENGTH
'0 0 20 40 M MAXIMINDEGRADATION, % wall FIGURE 6.2.3-3Collapse Pressure of Uniformly Thinned Tubing (from Reference 6 2-3) 6.,2- 22
BROKEN HOT LEG OF TUBE GUIDED BY UPPER PLATES TSP THK = 0.75
~ INCH TUBE SEVERED JUST 51.88 INCH BELOW THE FIRST PLATE TS Figure 6.2.3-4 Schematic of a Partial Tube
7.0 FUTURE ACTIVITIES 7.1 Metallur ical Examination As described in Section 5.2.3, additional tube sections were removed from the B-Steam Generator during, the Phase II repairs.
Many of these tube sections have been selected for metallurgical examination. This work is presently in progress at Nestinghouse and Battelle Columbus. The results of these examinations are scheduled for completion about May 1, 1982. These results will be submitted in an addendum to this report in early May.
7.1-1
0 7.2 Testin 7..1 ~t' As stated in Section 4.7, collapse and fatigue testing is presently in progress at Westinghouse. This testing is scheduled for completion about May 1, 1982. The test results and conclusions will be submitted in an addendum to this report in early May.
7.2.2 Combustion En ineerin Combustion Engineering is presently conducting testing relative to the effects of axial load and loose part impacting on a steam generator tube's propensity for local buckling. The testing is being performed using a model which simulates a single steam generator tube between the tube sheet and first support plate.
The test model is capable of applying axial load, peening,,and external pressure to the tube. Tube specimens will be eddy current and metallurgically examined following each test cycle.
The tests .are scheduled for completion in early May.
7.2-1
7.3 Loose Parts Monitorin S stem 7.3.1 Summar A loose parts monitoring system .will be installed on both steam generators at Ginna prior to returning the plant to power operation.
The system is intended to provide, within the sensitivity of the instrumentation, indication of foreign objects in either the primary channel heads or secondary side of the tube sheet. The system will be a Westinghouse metal impact monitoring system.
The system will include sensors (accelerometers) located on the exterior surface of both the primary and secondary sides of the steam generators. The sensors will be located to allow detection of a loose part in either the primary channel head or on specific the secondary side of the tube sheet. Figure 7.1 shows the location of the sensors. The sensors are capable of detecting acoustic disturbances in the steam generators. The system will be capable of detecting a metallic loose part that weighs from 0.25 lb. to 30 lb. and impacts with a kinetic energy of 0.5 ft-lb on the inside surface of the steam generator within 3 feet of a sensor.
7.
Sensor signals are transmitted to, a preamplifier mounted near each sensor. The preamplifiers transmit, a signal by shielded cable to a signal processing/data output panel located outside the containment building. The signal processing/data output panel will provide audio and recorder outputs, alarm outputs, and data processing. The alarm output will interface with the plant computer which alarms in the control room. An electrical one line drawing of the system is shown in Figure 7.2. The system will provide automatic, continuous, The on line monitoring of the system will be tested steam generators for loose parts.
prior to being placed in service to determine background noise and system frequency signatures as well as the frequency spectrum and on-line sensitivity for metal impacts.
7 3 1
~
e 7.4 Intermediate Outa e An intermediate steam generator inspection outage is planned at no more than 120 effective full power days (EFPD) after return to power. During this outage eddy current, fiber optics and video, and visual inspections will be performed. The purpose of these inspections is to assure that the corrective actions taken to preclude further periphery tube O.D. defects have been successful.
e 7.4-1
HOT LEG COLD LEG 2
4 TIRE LAIR I
TUBE QUET S/6 LOOSE PART MONITCR SENSOR LOCATIONS FIGURE< 7.I REV, 4/N/bR 7.4-2
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APPENDIX A Metallur ical Examination of Ginna Steam Generator Tubes*
- This report was also submitted to the NRC by letter dated April 23, 1982.
Abstract Detailed microscopic examiriations were performed on a section of Row 42
- Column 55 (R42-C55) tubing taken from the hot leg side of the R. E.
Ginna "B" steam generator in .the region where the tube burst. Five neighboring previously plugged tubes were also examined. All of these Inconel 600 tubes displayed on their 0.0. surfaces one or more axially oriented flat zones which contained circumferential striations. Thirty-eight of the forty flat zones that were examined exhibited cold work (O.D.) surface layers, a feature which is consistent with a wear pro-cess. The burst in R42-C55 occurred at one of these flats where the wall thickness had been reduced from the nominal 0.050 in. to 0.008 in.
Cold work was not identified on the flat that burst. Fractography revealed a normal ductile tensile overload failure at the burst.
Fatigue cracking was identified as one mode of breakage of previously plugged tubes. Normal metallurgical properties were identified for the burst tube and for one plugged tube.
'165s:10
- 1. 0 Introduction On January 25, 1982, a tube ruptured in the "8" nuclear steam generator of the Rochester Gas and Electric Corporation (RGE) R. E. Ginna Power Plant.
The burst tube was determined to be three tube-rows in from the periphery on the hot leg side at Row 42 - Column 55 (R42 C55).
turee The leak was due to an axially oriented "fish-mouth" opening just above the top of the tubesheet. In addition, some of the previously plugged neighboring peripheral tubes exhibited collapse, deformation, or. frac-conditions which were not present at the time these tubes were plugged.
A hole was cut in the shell at Column 55 and nine-inch lengths of the leaking and neighboring tubes were removed. Sections of six tubes were sent to the Westinghouse R&D Center, Pittsburgh; Pennsylvania for metal-lurgical characterization. These consisted of the leaking tube (R42 455) and five neighboring and previously plugged tubes (R44 .C54, R43 C54, R44 C55, R43 C55 and R43 C56).
2.0 Nondestructive Examinations Prior to removal of each section from the steam generator, a yellow dot was placed on each section to define the tube surface closest to the perimeter of the generator. In the current examination this dot was taken as the Oo position. Each tube length was received in one or two sections, originally extending from near the top of the tubesheet to ten in. above the top of the tubesheet.
Photographs of the leaking tube from Row 42 - Column 55 (R42:C55) show a "fish-mouth" opening at 0o and axially oriented flats at 0, 60 and 315o. These flats had, circumferential striations, Figures 2-1 and 2-2. Photographs in Figures 2-3 to 2-8 are of the neighboring pre-viously plugged tubes: R43 C55, R44 C55, R43 C54, R44-'C54 and R43 C56.
Similar flats were observed on all tubes at various angular positions.
There is extensive metal loss and apparent deformation on all tube sections. Tubes. from R44 C55 and R44 C54 arrived in two sections.
k Wall thickness measurements were consistent with a nominal 50 mil wall away from the flats and a wall thickness reduction at the flats, Figure 2-9, Table 2-1.
Double wall x-ray radiogr aphs were made of all tube sections at 0, 45, 90 and 315o. A Seifert Industrial X-ray unit was used with the fol-lowing settings: 5 min. exposure time,'110 to 70 kV, 10 milliamperes, 60 in. source-to-film distance, large focal spot, and Kodak M-8 lead pack film. Wall thickness reductions were easily seen; however, there was no evidence of intergranular attack or stress corrosion cracking.
0165s:10
On tube R44 C55, an EDS analysis on a flat was rich in Al, Si, Cr, Fe, Cu and Zn as compared to Inconel 600, Figure 4-5. The fracture surfaces near the bottom of the tube were not well defined and contained deposits, Figures 4-6 to 4-9.
Other fractographs on R44 C54 indicated that fatigue contributed to the fracture, Figures 4-10 and 4-11.
5.0 Com osition and Mechanical Pro erties of, Tubes An EDS analysis of Section 1A of tube R44 C55 indicated that the pr in-cipal elements were within specification. The mechanical properties of the Inconel 600 tubing were estimated and tabulated in Table 5-1. Knoop (500 g) hardness readings were made mid-wall on the metallographic mounts 4 inches down on Tubes R42 C55 and R44 C55. These were converted to Rockwell "B", RB, readings which were then used to estimate yield strength by two methods. One was a correlation established previously on Inconel 600 tubing, Figure 5-2, and the other was an International Nickel Company correlation on Inconel 600 sheet and strip. Reasonable agreement existed between the two conversions: all estimates of yield strength were between 45,000 and 62,000 psi. An estimate of ductility of the Inconel tubing was obtained with bends. Rings 7A, 7B and 7C from Tube R44: C55 and 7A and 7B from tube R42-C55 were used. Each ring was positioned such that a location free of flats between 145o and 225o would be at the apex of a U-bend. The bend was made by placing the ring over a 3/32 in. diameter mandrel with a line of contact at the desig-nated position and deforming the ring at this location around the bar.
This strained the O.D. circumference in tension. No fissures were observed on the O.D. surface at the apex and the calculated outer fiber strain was twenty-eight percent.
6.0 Conclusions
- 1. The microstructure and mechanical properties of the burst tube and a neighboring plugged tube were determined and were normal for the mill annealed Inconel 600 tubing material.
- 2. There was no evidence of stress-corrosion cracking or intergranular attack on I.D. or 0.0. surfaces of any tube.
- 3. All tubes displayed flat zones of O.D. wall reduction with circum-ferential striations within the flats.
- 4. Thirty-eight of the forty flats that were studied showed cold work on the O.D. surface, indicating that a wear process produced the
'lats.
- 5. The "fish-mouth" burst occurred at a flat where the wall thickness had been worn to 0.008 in. from the nominal 0.050 in. original wall; cold work was not identified on this flat.
0165s:10 J
5 qt
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2-1. Fish mouth crack at near Tuhe R42-C55 from RGE (GINNA), SGB, hot leg RM-94656
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4) 270'25'80'0'5' 00 Fig. 2-3. Tube R43-C55 from RGE (GINNA), SGB, hot leg RM-94658
Ptfigga
~e 315 Fig. 2-5 1800 Tube R44-C55 from 90'0 RGE (GINNA), SGB, hot leg RM-94660
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~<g 2-7 ~ Tube R44-C54 from RGE (GINNA), SGB, hot leg RH-94662
53.45>32,23,37 55 55 Pe e 56>44,26,17 23>44>55>56
~ % ~
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56 56>46$ 23 ~ 14-13$ 29>49 57 56 54,44,25,13-15,18,40,56 ~
58
' 56
~ C pig 2 9 Wall thickness measurements (mils) superimposed on 0'hotograph of R42-C55 RM-94664
7C 7B&ech.
SEM - Fracture face l - Surface etudjed lu metallography SEM - Fracture face Print of double wall X-ray radiograph at 0'or Tube R42-C55 and diagram showing cuts, section identification, and sample use.
RM-94665
7C 7g- mech.
7A. mech.
EDS 6
3A n I co A
3 2C Fracture surface shoMn in 135'hotograph 2B 'used in SEM study 2A 1A eH I C4 Chemical Analyses I
04 Fractography Fi 2 12 print of doub1e wall X-ray radiograph at top portion of Tube R44-C55 and 0'or cutting diagram RM-94667
R42
~
ct~~+
jog P~ R43 K
4 R44 C54 C55 C56 Fig. 3-2. Transverse cross-sections at ~ 4" from the tops of the sections RH-94669
Pig. 3-4.-'etallography on transverse cross-section M" towards tubesheet on section from Tube R42-C55
.4
+gal p>
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3-6. OD surfaces 4" down from the top of the section from Tube R43-C55 m-94673
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~ 8 Fig. 3-8. Metallography on transverse cross-section down 6" from top of top portion of Tube R44-C5S. Similar results were observed after polishing an additional 24 mils.
RM-94675
h) h Ii'PAh,+)P Fig. 3-lo. Continuation of metallography 7" from top of top portion of Tube 344-C55 RM-94677
~
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Fig. 3-12. 2 Tube R43-C54 RH-94679
Fig. 3-14. Transverse cross-section 22" down from the top of the top portion of Tube R44-C54 RM-94681
r 5
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Tube R43-C56 RM-94683
Table 3-1 Depth of cold work to the nearest 0.1 mil on wall reduced areas (Tabulated values are angular position-max. depth in mils)
Tube No.
(Approximate Axial Position From Top of Tube Section)
R4'.2-C55 2-1/2" '(0' 0.0)* (60' 0-4)*
4 II (0' 0.0)* (60' 0.4)
R43-C55 2-1/2" (30'-1.3 ) (225'-0.5) (270 -0.4) (330'-0.2) 4ll (30'-0. 3) (45'-0. 5) (225'-0. 5) (300'-0.4)
R44-C55 2-1/2" (60 -P. 5) (225'-0. 6) (270'-0. 3) 4 II (7 5'-0. 2) * (105'-0. 2) *(150'-1. 0) (200'-0. 6) (270'- 0. 4)
(105'-0. 5) (150'-0.8) *(270'-0. 6)*
7 ll (90:0 0} (180 -0, 5) ( 240 -0, 6) *
~
R43-C54 2-1/2" (30-0. 9)
- 4lt (30 0.9)*(195'-0.5) (315'-0.5)
R44-C54 2-1/2" (60'-0.9)*(90-1.0)*(270'-0. 6 ) (300'-0.9) 4 II (60'-l.l) A(105'-1.0)*(180 -0.6) (270'-0.7)
R43-C56 2<<1/2" (04 0.3)*
4lt (0 -0.5)*
- Shown in photomicrographs on transverse cross-sections
Table 3-2 Knoop (100g) microhardness traverses at various locations on Tubes R42-C55 and R44-C55 Tube Position No. Remarks > Depth from OD surface in mils 1 2 4 9 14 19 R42-C55 1 Flat 224 228 231 237 2 Flat 180 194 195 192 3 No Flat 258 224 229 219 R44-C55 4 Flat 321 300 223 '24 254 5 No Flat 364 261 312 289 261 209 6 Flat 321 312 248 214 226 194
Center of fracture OD
>r7 y I/
pP r,'lat on OD Fracture surface Pig. 4-2 SEM's of flat and fracture surface at center of fish mouth crack on Tube R42-C55 RH-94687
OD surface aeer peek in prior fissure Ar ~
P -
j'AQPFP
~ i J1 t
h
~
y her, vk pf 4 4 SEH' near peak fn prior ffgure and on fracture face.
RM-94689
Cut surface SI58il:
OII IgpW~
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~
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>135 Axial 135'135'xial Fracture Fracture Fig. 4-6. Fractography was performed on two axial fracture surfaces 1-1/2" from the bottom of the top portion of Tube R44-C55 JR-94691
e(
Flat <135'racture hxial Fracture Surface at
~ I~
Fig. 4-8 Fractography on <135'xial fracture RM-94693
a e e
'0 i
I Bottom
~ ~
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'D
- i ~ matching eurfacca Fig. 4-10. 180'iew of ring at bottom of Tube R44-C54 (Top) and fractography on bottom fracture surface showing areas studied in more detail RH-94695
F holllllllllslll Ills>>e(lll<h>
~Ally~sxuNlnllIlllIniNlIllllIIlllllNIIIlIlllllIllIlIlllIlIIllIllllllllllilIIllllllllllllllllull l<>>i>>NIIII lli>>>>>>>>>>>>>>>>>>>>>>>>>>>>>>>>~>>>>>>
EDS Analysis ASTM Spec.
SB-163 Element w/o w/o 6 Line Cr Ka 17.0 14.0 to 17.0 Fe Ku 8.5 6.0 to 10.0 Ni Kn 74.4 72.0 (min)
Fig. 5<<1. EDS spectrum and analysis for principal elements on Section 1A from Tube R44-C55 (Included are ASTM specifications) .
70 I
~
X
~
e~
u 60 0
~ 55 0 50 0 0
0- 0 8o,0
~ CI) 45 0 C)
+O Slope: 1,423 psi/RB o 40 0
35 72 74 76 78 80 82 84 86 88 90 92 Rockwell "B" Hardness Fig. 5-2. Correlation of Rockwell "B" hardness and yield strength on 24 heats of Inconel 600 tubing and estimated strengths of two Ginna tubes (dashed lines) .
APPENDIX B EPRI CE ANALYSES AND TESTS STEAM GENERATOR TUBING STRUCTURAL ANALYSIS B. 1 ~Summar The initial analytical efforts in evaluating the Ginna tube damage have been concerned with identifying potential sources of tube stresses. Emphasis in this work phase has been on flow induced vibrations and those factors affecting tube vibrations; This work was performed to provide general information and trends.
Analysis assumptions do not, in all cases, conform exactly to Ginna conditions.
A multiple span single tube was developed to assess tube vibration due to flow conditions. The model specifically addresses the influence of a compressive axial load on the tube frequency and the associated critical flow velocity. Results of the initial investigation confirm that compressive tube loadings produce a reduction in the tube frequency, and consequently, a reduction in the critical velocity to initiate fluid elastic coupling.
A multiple tube/support plate model was developed to define the variation in tube loads in the vicinity of a "hard spot" wedge location. Tube loads were developed as a function of the thermal expansion mismatch between the. shell and tubes. The initial analysis was performe'd assuming all active (no plugged) tubes and demonstrated that the tubes nearest the "hard spot" experienced the highest axial loads. 'Tube loads diminished rapidly with increasing distance from the wedge location because of the flexibility of the support plate.
The effect of tube plugging was also considered in the multiple tube model assuming that tubes were plugged subsequent to "locking-in" all of the tubes at 100% full power conditions. The results, after the unit is returned to full power, indicate tensile loads in the plugged tubes and compressive loads in the adjacent active tubes.
Considerably higher compressive tube loads can be predicted for the hot standby steady state condition; however, the associated flow is not sufficient to be of concern. The large compressive loads are not particularly damaging for only an infrequent occurrence of this condition.
The results of the work performed indicate that the plugged peripheral tubes at the "hard spot" experience a predominately tensile load during power operation (assuming that, the tubes were "locked-in" at full power and subsequently plugged). The absence of a significant compressive load in the tubes suggests that flow induced vibration is not the prime initiator of the initial B.l-l
damage in the plugged tubes. Some other initiating mechanism appears likely. However, once the initial failure is established, flow induced vibration can certainly contribute in the progressive damage.
B.2 Failure Se ence H othesis Loose part(s) exist in downcomer with sufficient downcomer flow forces to cause impacting of the loose part on heat transfer tubes.
B.2.2 Impacting of the loose part on peripheral tubes causes surface damage on the exposed face of the tubes.
a) Surface damage is generally of a peening nature, with possible metal loss.
b) Some I.D. surface disturbance may be present due to the impacting.
Work hardening of metal grains on O.D. surface will exist.
d) Free iron pick-up on the O.D. tube surface may exist.
e) Local dings may occur due to particularly heavy impacts which distort the tube through the entire tube wall thickness.
B.2.3
'Distortion of the tube from its original circular shape occurs due to force of impacts and/or due to bowing of the tube under peening action and bending moment caused by resistance to bowing at the support plate elevation. The distorted shape would be generally, but not necessarily exactly, in an oval foim with the long axis of the oval generally parallel to the tube bundle p'eriphery. The distorted shape would begin only some distance above the tube .sheet, since the tube is restrained to a circular shape by the tube sheet hole.
B 2.4 ~.Case 1 The tube is plugged due to false ECT indications of I.D. tube wall defects. False indications are caused by:
a) Vector additions in the ECT test of several forms of tube wall distortion, or B.1-2
b) A combination of work hardening of the O.D. surface changing the tube conductivity and/or free iron pick-up on the tube O.D. surface, or c) Copper deposits on the tube O.D.
ECT indications of a "bulge" are caused by the distorted tube shapes mentioned in three above.
5.2.5 ~Case 2 The tube is plugged due to I.D. defects caused by coriou cracking.
Coriou cracking results from high tensile strains at the tube I.D. due to some combination of:
a) An axial tensile load in the tube during transients, caused by locking of the tube in the support plate in the hot condition. Locking may be in part due to the reaction to bending loads caused by bowing of the tube under the peening action.
b) High I.D. tensile strains existing at local dings in the tube wall caused by particularly heavy impacts (see B.2.2e).
c) Other tensile stresses resulting from internal pressure, tube vibration, distorted shape, etc.
(Note that, if tubes become locked into the support plate at power, then compressive loads are developed in active tubes) full B.2.6 Tubes are plugged preferentially at welded lug locations because of hydraulic pertubations in the downcomer flow from upstream effects (flow eddy created below connection to the shell) and downstream effects (absence of flow holes in the support plate changes the flow drawn into the tube bundle at these locations).
These hydraulic pertubations cause the loose 'part to dwell at these locations for most of the time rather than continue an otherwise random movement around the downcomer annulus.
B.2.7 Once plugged, tubes without a breach in the primary-secondary boundary are under external pressure during power operation (approximately 770 psi at full load and 970 psi at hot standby).
B.2.8 I Plugged peripheral tubes experience combination of:
local collapse due to some B. 1>>3
a) External pressure.
b) Local bending moments, tube ovality, or deformation (Items 2.e. and 3. above).
c) Axial tensile load (Item B.2.5b).
d) Particularly heavy impacts from the loose part occuring after the tube is plugged.
B.2.9 Flow vibration effects on the locally collapsed tube result in fatigue failure of the tube in the collapse region.
B.2.10 Fatigue failures progress to the point of parting the tube, allowing that portion of the tube above the failure elevation to behave as a contelivered tube under flow vibration forces.
The failed tube acts at a "loose part", subjecting itself and neighboring tubes to impact and wear damage. Neighboring tubes are subsequently plugged due to ECT indications resulting from the damage.
B.2.12 Progressive tube plugging, tube failure, and damage to additional neighboring tubes occurs, resulting in conditions observed during the current outage. The tube wall in R42 C55 had likely experienced considerable metal loss by wear or impacting prior to failure, creating conditions which result in a rupture of the length observed.
B.3 3-D-Plate and Tube Anal sis B.3.1 Statement of Problem The purpose of this study was to determine the loads on the steam generator tubes as a result of the tubes becoming locked into the support plate.
B.3.2 Summar and Conclusions The loads considered in this analysis were the result of a thermal expansion mismatch between the tubes and the stub barrel, and the thermal expansion mismatch between the hot tubes and the cold (plugged) tubes.
1 The tubes were considered locked in hot (100% steady state) condition. The first condition analyzed was shutdown condition
(70'F) prior to certain tubes heing plugged. The tubes closer to the hard spots were found to be under most tensile loads. The second condition analyzed was 100% steady state condition after a certain number of tubes were plugged. The cold (plugged) tubes were found to be under tension whereas hot (active) tubes surrounding the cold tubes were under compression.
B.3.3 Results The results of this analysis is presented in graphical form in Figures B.1 and B.2.
Figure B.1 presents the worst loads for the shutdown (70'F) condition prior to any tube plugging. The loads were as a result of initial hT = 72'F between the tubes and the stub barrel.
Figure B.2 presents the worst loads for the 100% operating steady state condition
' after certain tubes were plugged (TS~<< = 483'F,
TUBE Pr.UGGED TUBE B.4 Sin le Tube Model for Vibration and Fluid-Elastic Instabzlxt Stud B.4.1 Statement of Problem r The purpose of this study is to evaluate a multi-supported single tube loaded in compression to determine its vibration characteristics and susceptibility to flow induced fluid-elastic vibrations.
B.4. 2 ~summar The single tube unloaded configuration has a first mode natural frequency of vibration of 24 HZ. If the tube is assumed to be locked in at support plate elevations 1, 3, and 5, during appli-cation of a uniformly distributed axial load, buckling occurs in one of the upper spans at a force of 1360 lbs. If assumed to be locked-in at the first suport elevation only, the tube is buckling occurs at 3340 lbs.
The axial loads necessary for initiation of fluid-elastic instability are, 1300 lbs. and 3220 lbs. respectively, for the two conditions discussed above. For these loads, the critical velocity is equal to gap velocity, 11.5 feet per second.
B.4.3 Results Figure B.3 shows the variation in natural frequency with compressive load in the tube. Curve A was obtained by restraining vertical movement, at support plate (S.P.) 41. All supports restrain lateral motion. Buckling occurs in the first span above the tube sheet for an axial load of 3340 lbs. Curve B was obtained by restraining vertical motion at SP-1, 3 and 5. A uniform axial load was applied in all spans from SP-5 downward. Buckling occurs in one of the upper spans at 1360 lbs. (Frequ'ency = 0 at
I
. buckling load). Figure B 4.shows the variation in critical flow velocity with compressive load. Curves A and B were generated using the frequency-load results presented in Figure B.3.
For Curv'e A of Figure B.4, S.P. 01 locked, the axial compressive load required for initiation of fluid-elastic instability is about 3200 lbs. at 11.5 feet per second. Curve B with S.P. 011 3, & 5 locked shows a required axial load of only 1300 lbs.
B.5 Thermal Anal sis of Secondar Shell-Tube Sheet Juncture Re won Hot Le Sxde B. 5. 1 Statement of Problem This calculation presents a thermal analysis of the secondary shell-tube sheet, juncture region for the Ginna steam generator for the purpose of comparing mean shell temperatures with secondary tube temperatures for use in a stress analysis. Several operating conditions are investigated.
B.5.2 Summar and Conclusions A brief summary of the pertinent results of the analysis is presented below including the mean shell temperatures and tube temperatures for comparison. The end points of the 100'F/hr transients are presented here:
T T T m m m Transient Shell Active Plugged Remarks Tube Tube
- a. Plant Heatup 529oF 5450F 5450F End pf 70oF tp 5450F Ramp at 4.75 hr.
- b. 0'/ Load S.S. F 5450F 545oF Isothermal at 545oF
- c. Plant Cooldovn 86 F 70 F 70 F End of 545 F tp 70 F Ramp at 4.75 hr.
- d. Cold Feed at Hot Standby S.S. 150 F 530 F 131 F 200 GPM Flow at 40 F
- e. 100'/ Load S.S. 475 F 555 F 475 F CR = 3.75; 81'/ of F.W. to Hot Side
- f. 100/ Load S.S. 485 F 555 F 484 F CR = 4.70; 81'/ of F.W. to Hot Side B.5.3 Results Computer output showing representative temperature distributions for the various loading conditions and from which the summary of results in Section B.5.2 was obtained is presented in Figure B.5, sheets 1 through 9. Figure B.6 shows the computer print pattern corresponding to the element layout of Figure B.7. Detailed B.l-6
temperature vs. time curves, including the outermost tube temperatures, are presented in Figure B.8, sheets 1 through 3.
B.6 Thermal Anal sis of Downcomer Annulus B.6.1 Statement of Problem Determination of stresses in both active and plugged tubes is part of the investigation into the damage in the Ginna steam generator. The temperature of the downcomer annulus flow during steady state and transient conditions is needed to establish shell and tube temperatures for these stress calculations.
B.6.2 Summar of Conditions The analysis was performed to evaluate the temperature of flow in the downcomer annulus for the conditions analyzed. The results were used directly to establish tube temperatures and indirectly by providing input temperatures for the secondary shell - tube sheet analysis. During hot standby at 544.6'F (1000 psia) without recirculation, 40'F feedwater to the hot side is heated to 130.7'F.
At 100% power, the temperature of the mixed feedwater and recirculating saturated water increased only 0.2'F.
B.6.3 Results The temperature distributions for the three conditions analyzed are presented in Figures B.9 through B.ll. The results of the
'transient downcomer outlet temperatures vs. time are plotted in Figure B.12.
B.7 Tube Wall Tem eratures B.7.1 Statement of Problem Average tube wall temperatures for the length between the tube sheet and the support plate are needed to calculate stresses due to axial loads.
B.7. 2 Summar and Conclusions The average temperature of plugged tubes is the same as that of the surrounding fluid, while active tubes and plugged tubes are the same temperature for heatup, 0% load steady state, and cooldown.
At 100% power, there is boiling - either sub-cooled or saturated-from the tube sheet to the first. support, the length of interest.
The computer program SGTUBE shows an average heat flux of 116,260 Btu/hr-ft~ for this length. The average wall temperature for this value of heat flux is 555'F.
In cold feed at hot standby, the average tube wall temperature is calculated to be 529.7'F. The primary temperature is at the secondary saturation temperature, and heat transfer to the sub-cooled secondary is by natural convection.
B.7.3 Results sec, F bT, F t ~tf-Flux L.M. Flux Btu r-ft~ wall, F 130.8 378.16 59,466 1926.6 50,644 19.67 525.4 250 268.98 42,741 1601.6 34,356 24.10 531.6 350 178.32 27, 146 1710.4 18,206 48.57 534.3{1) 450 87.70 11,485 The wall temperature at T = 350'F is extrapolated for the 8.04" to the support platI.
B.8 Feedwater Distribution to Hot and Cold Le Side Downcomer B.8.1 Statement of Problem The feedwater rings on the Ginna steam generators have been modified to offset the feedwater flow to achieve a 80%/20% flow split to the hot leg/cold leg sides of the downcomer. Also, J-tubes have been installed on the top of the feedwater ring to reduce the probability of occurrence of water hammer. The purpose of this analysis is to calculate the feedwater distribution for 100% power operation to verify reported values and to calculate the flow distribution for hot standby with 200 GPM feedwater flow.
B.8.2 Summar and Conclusions Results of the calculation for the modified feedring with J-tubes show that 81% of the feedwater is distributed to the hot side and 19% to the cold side for 100% power conditions. This confirms the prediction that the J-tube modification does not change the previously achieved flow split. The flow split for 200 GPM feed flow at hot standby was calculated to be 80%/20% to the hot leg/cold leg sides.
B.8.3 Results The following is a tabulation of results from RINGFKO. The flow shown is one half the total since by symmetry, only one half the feedwater ring was used in the model.
B.1-8
tl 0
100% Power Hot Standby Nozzle Flow X Flow X No. lb/hr % Flow % Flow lb/hr % Flow % Flow Vent 63915 4.65 2506 4.99 1 64393 4.68 2462 4.90 2 65614 4.77 19.0 2480 4.94 19.8 3 66800 4.86 2499 4.98 67835 4.93 2509 '.00 5 69102 5.03 2542 5.06 6 70296 5.11 2574 5.13 7 71415 5.19 2eo5 5.19 8' 72457. 5.27 2633 5.24 73419 5.34 2659 5.30 10 74298 5.40 2684 5.34 11 75092 5.46 81.0 2706 5.39 80.2 12 75800 5.51 2726 5.43 13 76418 5.56 2743 5.46 14 76946 5.eo 2758 5.49 15 77381 5.63 2771 5.52 16 77723 5.65 ,2781 5.54 17 77968 5.67 2788 5.55 18 78067 5.68 2791 ~ 5 56
~
B.9 Downcomer Flow Velocities B. 9. 1 Problem Tube damage in the steam generator may be caused by vibration and by impaction from heavy objects. Both of these causes are dependent on the velocity of the downcomer flow.
B.9.2 Summar and Conclusions The velocity of the circulating flow exiting the downcomer is calculated to range from 11.08 to 16.35 feet per second, depending on the secondary steam flow specified and the circulation ratio.
These give velocity heads or impact. pressures of .66 to 1.38 psi. As the net weight per square inch of 7/16 inch thick iron or steel is .112 pounds, turbulence and deflected flows could easily lift such as plate.
The approach and gap velocities as calculated are an average based on the opening below the wrapper. Actual velocities will vary axially, due to the jetting action of the fluid leaving the downcomer.
B.1-9
During hot standby, there is a flow of 200 GFM of 40'F feedwater.
Assuming no recirculation, the downcomer velocity is .086 feet per second, and approach and gap velocities are .013 and .044 feet per second, respectively.
B.10 Steam Generator Thermal-H draulic Anal sis B.10.1 Statement of Problem The thermal-hydraulic characteristics of the Westinghouse U-tube vertical steam generator were calculated in order to understand various phenomena observed during inspection of the R.E. Ginna nuclear steam generator.
B.10.2 Summar and Conclusions The vector components of the secondary fluid velocities, the quality, and the vapor fraction of the two-phase mixture near the tube sheet and through the first tube support device were calculated in detail to determine the contribution of fluid conditions to the observed phenomena. No unusual trends were observed. However, the analysis and extensions of it may be useful in understanding contributions due to the loose objects discovered in the steam generator.
B.10.3 Results The radial and axial velocities of the secondary fluid, the quality, and the void fraction at select locations in the steam generator are presented in the following line printer plots (Figure B.13 Figure B.26). The key for interpreting these plots is shown at the top of each figure. Each plot integer spans 10% of the range of the variable in that figure. Even integers are omitted to enhance clarity.
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BASS QUALITY ~ ~
AT THE HORI?ONraf,. PL.ANE IZ"-.
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,7.. 999:.."'99 ) '. '"'5qgqSg) a a a T SIDE . COLD SIDE equality at 7.in above tubesheet (CR = 3.76)
Note: 'owncomer opening is 14 in high.
FIGURE B. 13
RADIAL VELUCITY (HI X TUKE) AT T](E HOB I7QNTAL PLANE IZ 0
RANGE UP If lE F I ELO IS ~ SPO TO ~ 44'5 meter/sec
- ".'!
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Note: Downcomer opening is l4 in high.
't' FIGURE B.14
~ '
AXIAL VELOCI TV (Nf<TUHE) AT .THE HORI7DNTAL PLANh: lZ-"" 1 R+HGE UF THE F I ELf) Is -. 2nD T n 1 ~ 22 meter/sec C
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HOT SIDE COLD SIDE Axial velocity (m/sec) at 14 in above tubesheet (CR = 3.76)
Note: Downcomer opening is 14 in high.
FIGURE B.l5
~ ' e see I ~
1 D CTIUN (RG) AT THE HD"17DNTAL PLANE IZe FIELD IS n. Qxo RANKLE OP
,8"ND SCALES ARE Tf1E 1
5
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t
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e e a<< area<< ~ I 5 Q~)9 7 / 5 ] Qeeaeeveeeeeeeeeeeeae 999999 / 3 1 f1 e ee a eeeaae eve eeeeaeeeaeeeeee ~ eeeseassepaeeeeeeeee ~ eeevaeaeeeeeeeeeaeaeae ROT SIDE COLD SIDE Void fraction at 7 s,n above tubesheet (CR 3.76)
Note: Downcomer opening is 14 in high.
FXGURE B.l6
~
I 3
.'thy',>aglTV ar THE HDRI70DttAL PLaWE l7= 2 RA."t;E liF THE FIELD IS 3n2F<<0] TO .3p5E-n]
-.>Etc'-n] ]U .2/SF']
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3
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FIGURE B.17
l ~
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940 1 AL VELUC L T'( (~t~tut F.) AV THE t<O~I70Wr4L PL4NE IZ= 2 RAnr;E u; ft>E FIELO lS 275 ~ 276 meter/sec 9Hw() $ C ALES 4HE 1 . TU <<,220
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55 777 7 553 J
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~ a <<a a a <<<< ~ a ~ w ~ w ~ ~ a ~ <<<<a ~ ~ ~ ~ <<~ w a a <<<<m ~ ~ ~ w < ~ <<a a a m ~ ~ ~ m ~ a <<a <<m ~ <<a <<
g HOT SIDE COCO SjOE Radial velocity (m/sec) at 20 in above tubesheet (CR = 3.76}
FIGURE B.18
~ '1 XIAL VELOGI I Y (HI>TUBE) AT THE HOR I70N T AL. PLANE I 7." 2 RAWGE DF THE FIELD I~ .2p9 T<f 1 ~ a9 meter/sec
~l<O SCAAAb h~F. 1 id29 rG=.>svh-ni p 3 i l c15 TO, .'up7
~ 6f9 TD ,831 7 1 i p<> Tu 1,25 9 1 ~ 47 aa aaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaaa aaaaaaaaaaaaaaaaaaaaaaa I ~ ~ 0 '
~ ~ ~ ~ t~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ '
I
~ P ~ ~ ~ ~ ~ ~ 0 f ~
1
~ ~ + ~ ~ ~ ~ ~ I I~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 4 ~ ~ y ~ S'33%33 '
~ e ~ ~ ~ ~ ~ o ~ ~ ~ ~ ~ ~ ~ I I ~ ~ ~ ~ ~ ~ e e ~ ~ ~ ~ ~ 53533 35z3 5355 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ I J ~ ~ ~ t~ ~ ~ ~ 'i ~ ~ 333333353533 cS3335333 ~
y ~ ~ ~ ~ o ~ ~ AT
~ ~ ~ ~ ~ ~ ~ ~ ~ 33 X3 55 333333-S3 Z3q3.353333~ ~..... ~ ~ oI J, ~ ~ ~ ~ ~ ~ ~ 35 35+3 S 5 3S353515333S 3533 z3 ~ ~ ~ ~ ~ ~ ~ I I~ ~ ~ ~ ~ ~ 3'S.5553333 '3533.S33333 ~ ~ ~ ~ 'e ~ I I ~ ~ ~ ~ ~ 3353533 33S3<3333 ~ ~ ~ I 1 ~ ~ ~ ~ 555z555')5 353 33z35355 ~ ~ ~ 0 I I ~ ~ ~ 55'55 55 33333333 ~ ~ ol I' S5c) 777777 5 3S.S 1
1111 ~
3S33533 ~ ~ I Ii 555 5'55 Tl. 7 7 55 111111. 555313
~ 'I ~ S 7 999 7 5~ 3S 1111111 335333 3 555 7 9 999 77 3 11lfl1111 333353 7 9 i>99 7 5 1111111111 97 5S 1 1 11111111 f 1 cj 7 1111 aaaaaaaaaaaa ~ aaa <<aa aaaaaaaa>> a aa a aaaaaa aaaaa a ~ aaaaa HOT SIDE COLD SIDE Axial velocity (m/sec) at 27 in above tubesheet (CR = 3.76)
FIGURE B.19
4 ~ ~
~45" uaaLlfr az r>IE HURI708]4'L PLANE IZ= 3 waif' QA)ICE t)E F IELP IS 51 ]F 0] rii ~ <<q5~-nl Sent.p 5<ALES'ARE -.3))c-nf TO -,240E-n]
3 - ]7uF.-n] TO .99qE-A2 5 ".890F.-n2 Tt]
,( 7 E - nl T (I . .
.010E"nE.'1IE
~ 1 SEE-nI
.-' . +25SEan] TO '",325E "01,
'I .. ~
4 a aaaaaaaaaaaaaaaaaaaaaaa aaaaaaaaa <<aaaa aaaaaaa
~,'e
'a aaaaaaaaaa 0 t~ ~ ~ ~ ~ ~ '
~ ~ ~ ~ ~ ~ ~ t ~ ~ ~ 533355 1 ] 1 ] l ] e ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ g ~ ~ ~ ~ ~ ~ 1 I
~3"
1 I~ ~ ~ ~~~~ ~ ~ ~ ~ ~ ~ ~ a ~ 33333 3353S5 ] 1] 1] 111]le ~ ~ ~ ~ ~ ~ ~ a ~ ~ ~ ~ ~ ~
.:I ~ ~ ~ ~ . ~ Y33Y 1111111 I 111' ~ ~ ~ ~ ~ ~ ~
~ ~ i I
~ 'I
- ,'
- I ~ ~ ~ ~ ~ ~ ~ ~ ~ 3 555555 '3333 l 1 11 111 11 1 1 ~ ~ ~ ~ ~ ~ ~
'i) 1] ) ] ] ] ] ] ] 11]1..."
~ <
~ ~ ~ ~
.'.: I ~ ~ ~ ~ ~ ~ ~ e ~ 5555'35555'5')5 333 1 ~ ~ 0 ~ I I ~ ~ ~ e ~ ~ ~ 55~55 S5> 333 1111] ]1 111111 ~ ~ ~ ~ ~ ~ ~ I I~~ 5555, 55 33 ]'l]1 1]1111] o ~ I I
I
~ ~
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~ y ~ ~
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5>i5 S~(77777 777777 (777 555 SS 3'33
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))]ill))]
~ ~ ~ ~
~ '
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555 555 7777 7 7 7
. 7777 77 I 55 5'-3 333i333 333Z333 53 l))]}))))reI
- 11111111,,7
~
J ~ 777 99999 7 1]11]1]1 I ~
I. 555 77 99999999 17 55 5~35533 ]11))))]qI.
I 77 999~ 999 77 55 355353 11] I J. 99 / 35 S5733 I I 7 55555 353 I
~
aaaaaaaa 9 5 I
<<aaaaaaaaaaaaaaaaaaaaaaaaa aaa a<<aaaaaaaa>>aaa aaaaa HOT SIDE COLD SIDE Quality at 33 in above tubesheet (CR = 3.76)
FIGURE B.20
QAI) 1 AL V];],UC 1T Y (HiwTU<F) AT THE HU<<I7])<f AL P]-ANE IZ=
!IHHBEE UF IHHE FIELL'3 .IHHE ~ 113 . meter/sec SAW]J Sch].ES AwE 1 'E)2
" 'OTO ']62 5 131 ]01 5 -. ndF.-n] 10 -.3')7f-n]
7 .'92')E-n2 TO .2]?E-n,]
9 .5] bF -n] TO .82] f n]
a>> <<a<< a>>a aa <<a<< >>a <<a a<<a a<<a a a a a a a a a a >>a a a aaa <<a>>a<< a <<r a a a a a a a a>>
I>> ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ r ~7 77 ~ ~ ~ ~ ~ ~
'1 r ~ ~ ~ ~ ~ r>>
5
~ ~ ~ ~ ~ ~ I
~ ~ ~ >> ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 0 I 555 555 5b5b 3 55 I~ ~ ~ ~ ~ ~ >> ~ ~ ~ ~ ~ ~ ~ ~ ~ I I~ ~ ~ ~ ~ ~ >> ~ ,>> ~ ~ ~ ~ 1 55555 q55555 7.... .. ~ ~ ~ ~ ~
EH55 35533533333 5333 7. .. ~ ~ ~
I........
~ ~ ~ ~ ~
~ 53 53335333 . , 5353533 <3< 7 ' ~ >> ~ ~ ~ ~ >>?
T ~ ~ ~ ~ ~ ~ ~ 7 551 333333 1 1 1 1 ~ 33333 555 '7 ~ ~ ~ ~ ~ ~ >>I f >> 4 ~ >> s ~ 5 3$ 1 1 1 1 1 ] 1 ) 1 ] 1 '1 1 1 333 555 ~ ~ ~ ~ ~
I ~ ~ ~ >> ~ 5a 3.) 1 1] l)]11111]1] W3W 6b ~ ~ ~ >> I J>>>>>>>> 55 '1]111] ] ] ]1 .%35 555 ~ >> ~ ~ T 33<33383333 ))1]11) ) ]1] 1 333 q55 I I 5 33 3533 55 55~5555
'555
~33 333 33
]1)1111] ] l 1] 1
]111)1]i))11)]1
]
]11]1] 1111] 1111 ] 353 5
'3 33 555 55
~
5>>I' 7
~
~
>>I
~555
~
J 553>>3 777 77 11111111111111 11 1 .7, 5 7 777 5 3 1) 1111).l 1 ) 11]1] 1 999999 753 ]111]11),]- I
) ~ H 1] 11
~ <<<<<<a a <<a a a <<<<<<<<a a <<<<<<a <<a <<a a <<<<<<a a <<<<<<<<<<<<<<<<a <<<<a aaa HOT SIDE COLD SIDE Radial velocity (m/sec) at 33 in above tubesheet (CR = 3.76)
FIGURE B.21
l
~ l lASS uuAL 1 1 T AT THE hUR 17U~< 1 AL PLAt4E IZ"- 6 RA )QE UF Tt1P f'IELU IS 'q>K 01 To . ~ BPlE .1)1
.Vnnr;-nt r8 . ATe-n>
3 ~ 408E n1 T U,'o60E-pl 9 ~ 'nlLF-nl ln .-'563E-nl 7 .Bynp-nj Tu .6nhE-nl 9 .717E-.nl TU '.,768E-nl aaaaaaaaaaaaa aa a a a a a a a a a a a a a a a a a a a a aaaaaaaaaaaaaa aaaaaa
~ ~ ~ ~ ~ ~ ~ 5555555 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ fI I' ~ 0 0 ~ ~ ~ ~ ..-....5555 55SS5~55 ~ ~ ~ ~ ~ ~ ~ ~ ~ '
~ n ~ ~ ~ T
~ ~ ~ n ~ ~ ~ ~ ~ ~ . n . &SS'5'iS>~&55>55%
~ ~ ~ ~ ~ ~ ~ t~~~~~I I~~~~~~~~~ 555555 555555555S ~ ~ ~ g ~ ~ ~ ~ ~ n I
.I ~ ~ ~ ~ ~ ~ ~ ~ ~ S55 5555555 5~5zb ~ ~ ~ ~ . ~ ~ ~ I I~ ~ ~ ~ ~ ~ ~ ))')77/ 55'~ 55bqÃSs 5$ ~ ~ ~ ~ ~ ~ ~ I I ~ ~ ~ ~ ~ ~ 7 ~ 7? 7 7 ~77 12 555 555'i 7 7 ~ 5 '
~ ~ ~ ~ ~ I In' n777( 7777 '555 5 5'.i 7 SS ~ ~ ~ ~ I
~ ~ ~ ~ 5>~ t~~
'OLD
~ T I ~ ~ ~ 77 777 SSS 333%33333;4 5555 555 ~ ~ ~ T, I ~ ~ 77 999 777 ~5 333m 8333 5555%55 FAT 7 9999 >999099999 77 555 335 334 55%55 53 I J
ill lg ~
I ~ 9 9') 99q ) 9 ) 9~~999999 7 f 55 333 333 F55 3 ~ I I 999 9999 7 33 1111111 333 bS I 999 7 0 3> 111111111 3 .I I .99 55 3~ c 1 l l l 1 1 I T 99 11 I aaaaaaaaaaa a HOT SIDE SIDE equality at 91 in above tubesheet (CR = 3.?6)
FIGURE B.22
RaqlaL vELljt:LTV ( >I<TURF j .AT 'THE HURI7Qj<raL PLawE Iz=
RAwGE UF I f<E F I ELG I3 ". 5nO ~ 27~ meter/sec BnND SCagES aRE I ~ << ~ sou io -,pea 3 I>> LBc< To a;(c37
-.n!IVE~nIII=.I l1 c=nl
~ '>ATE A1 TU ~ I Oq 162 .TU ;220
<<a aa <<a a a <<a <<<<<<<<a a a <<<<a a a a <<a <<a a a a <<a <<a a a a aa <<a <<<<<<<<a a a a a <<a<<a aa<< I I ~ ~ ~ ~ ~ ~
~ ~ ~ ~ ~ ~ ,~ ~, ~ . ~ 5 S555555555555q5 i55555557S5555555555555
....................t
...,, ~,, ~..., I I~~~ ~ ~ ~ ~ ... ..~ 35555555 I~ ~ ~ ~ ~,~ ~ . ~ ~ . '5555565
. 55,55i55Y5Yi I 1
I I~
>> ~ ~ >> ~ ~ ~ ~
~ ~ ~ ~ ~ ~ ~
~ ~ ~ ~ ~
'55555555S555 55>5
.'3 333w3S33 55555m 55555S S555s
~ ~ ~ >> ~ ~ ~ ~ ~
.....I
>> ~ >> >> ~ ~ I
~ ~ ~ ~ ~ 777777/~5 5 3%3i333jj3 A&55.....I I~ ~ ~ ~ 777'7 7777 5555 3333%3q333333 555~5; ~ ~ ~ t Z ~ ~ ~ 777 777 555 ljaj<333353333> 55555...1 I, ~ ~ 777 ~IIVV 7~7 5 33333ij33333333%'j .5555 ~ ~ I
~ 7 77/ 9~~9'3999 7 55 3333'33w333333333335 55SS ~ I I~ 77 7 9~ 90'9ct9 77 35 333335:45Z3Z 33333353 555S>>I 5 33333'534533w '
c9 7 3335 c~ 7 5 33 3q3 I
<<a <<<<<<<<<<<<<<<<<<a <<<<<<a <<<<<<a a <<<<
HOT SIDE COLD-SNE Radial velocity (m/sec)'t 46 in above tubesheet (CR = 3.76) t/ote: The first tube support device is at 51.8 in above tubesheet.
FlGURE B.23
Ax]AL vELDCI fY (')] "Tu<H "T T))F. HOB]7()NTAL PLANE ZZ= 4
~ "'..., ..'
RANGE DF s~aL
)'HE 3
FIELD lS a<3 I 3 -,<72
.6~9
] ~2'5
].g3
.8abF.-O]
'TD
)'t)
TO TO
.27S
., <6t3 F 05 1 ~4ff 1 ~ 83 2, n2 meterlsec a a a m ~ ~ <<>> ~ aa m ~ ~ ~ ~ w ~ aa aa ~ m a w ~ a ~ ~ w ~ ~ ~ ~a~e a awaa~~aaae<+~a+~a'+<~
I ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ '
~ ~ ~ ~ ~ ~ )))])))] ~ ~ ~ ~ ~
111 1111111,.....
yy ~ ~ ~ t~~ ~ ~ ~ ~ ~ ~ ~ T
.. I T ~ ~ ~ ~ ~
I~o~~~o I
~ ~ ~ y ~ ~ ~ ~ ~
~ e ~ ~ ~ ~ ~
~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~
3 533333'S 335'5 45:5'3 53 333~3 <
I I r I I I I I I I 'I I
) ]) 11) ) 1,) ]11)
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~ .
. ~
a 1 1 >> ~ ~ ~ ~ ~ ~ ~ ~ ~ ~
~
I T ~ ~ ~ ~ ~ ~ ~ ~ +
I 333 3333 333 1]i 1) )1)1)
) ]l])1) ~ ~ ~ ~ ~ ~ ~ ~ I
~ ~ ~T~33 93339339 1111111111111111111, ~ . ~, I
])])))))))])])))))1)]
~
I ~ ~ ~ ~ ~ ~ 33 55'3555'5555555 33 ~ ~ ~ ~ ~ ~ T
? ~ ~ ~ ~ ~ 35 55 555> 53 ] ] 1']
) ) ) )))) ) ] ] ]
1 ] )) ) ] ] 9 ~ ~ ~ ~ I
) ~ ~ 3 0>'i 77777 1) 1) 1) ).11) 1) 1) 11] 1] ) 1] ] ~ ~ ~ ~ I I ~ ~ ~ 3 55 7~777777777 b0 3 111) 1] ] 1 ]1111)11111) 11] l e ~ ~ T
) ~ 93 55 77777 771 5 1) ]1])1)))1)1) ))11]111]1]o9T I~3 9r5 rr 77 77 7 9 I )))]1))]1))111)1))11)1]1))1 I I'555 777 999999 9o99 999 9 <r9
'77 '5 .
1])l)l)))])])])])1))1))'))]1I
~
~
777 777% 1) 1) ) ) ) 11) 11) 1]111) ]1 el
) >
9<9 99~)9 II '1111111111 I'I I 11
))])
99 777'I) ) ])11 T 99 7 111) ww%%%%%&~~~a~~ggwyg~aaa~~~~%~&w&%~&aaa&& 0 HOT'IDE COLD SIDE 7
Axial velocity (m/sec) at 51.8 in above tubesheet (CR = 3.76)
Note: The first tube support device is at 51.8 in above tubesheet.
FIGURE B.24
0 ee qhSS t77JALI f Y 47 THE HO~<I7UHTAL PLANE IZ-"4 "Al<GE UF 'HE FIELD IS . 2c,BE 01 TO .47OE-01 1'
E 119F nl TO 447 E 02 5 .49dF-n2 TO,104E-.nl 7 7755~0 TO P 53E 171 9 .328F.-nl TO - ~ 402E-01 I ~ ~ ~ '
e e e e ~ ~ ~ ~ ~ e ~ ~ ~ 3333333 1 1 ]1 1 1 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ e ~ ~ ~ ~ ~ ~ ~ ~
I~ .3333 11111 I
~ ~ ~ ~ ~ ~ ~ ~ ~ e ~ er ~ ~ ~
~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 5553 5 555555 )5555555>
>'llffDDl1 11 11 ~ ~ ~ ~ ~ ~ ~ ~ e ~ ~ ~ ~ ~ ~ ~ T
~ ~ ~ e ~ e ~ ~ ~ ~ ~ ~ ~ ?
I I
~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~
~ ,... ~ ~ ~ ~ 5555 . 555 353 333
) 1 > 1 1 1 ) 1
)11111111111111 a-f1 1
1~~1>
1 t 1 1
1 1
'fll
~ ~ ~ ~ ~
~ ~ ~
~ ~ e ~
~ ~
~
1 I~~ ~ ~ ~ 55 77/777777/7 555 35 111111111)1 I 11llllllll
~ ~ ~ ~ ~ ~
I~ ~ ~ ~ ~ 5 7777/77'/777'/777 55 33 ~ ~ ~ ~ ~ I
~ ~ ~ , 7/ A7 Ti7 lllli1111 I e ~
.777 77/
55955
'7 777 55 3q3q33 55 3333333.5 1111<111>
111111111 ~ ~ I
.I I
I
~ 77/ 9AMa99E) 99999 gg:~99 c)9q9
// 5 q5333333 333538
~ 1111 f1111 I 11111111 ~ I 1111 I
- eT7>~ 3333333 . 1 9999 3333 99 I aw ~ mw'eww ~~w~m~ww ~ ~wI~~~W~~Wammmm~a W~~Wmwwn~~~Wme~~mf W~mmw HOT 51OE COLD SIDE equality at 46 in above tubesheet (CR = 3.76)
Note: The first tube support device is at 51.8 in above tubesheet.
FIGURE B.25
I QA f) AL VtlijClr V ('1 I P > itaE) Ar r)<E HO~ I70N I AL PLANE l l.=
iAt'GC ~F fly'E F IELl) IS ~ Be5 met~er sec gr,vP SCALLS A~K 463 rg <<,'382 500 ro '218 137 to=,eszt'.-et
~ 2hwF."n I tQ ]05
.f90 ro ~ 211 aa<<aaaa'<<aa'<<a'aaaaaaaaaaaa'aaaaaaa<<aa a a a a a aaaaaaaaaa'a'aaa I~ ~~ ~~~~~~ ~ ~~ ~ ~ ~ ~ ~ e ~ ~
7 I 17717771 ~ ~ ~ ~ ~ ~ ~ ~
~ ~ ~ e ~ << ~ ~ ~ e ~ I m
I ~ ~ ~ ~ ~ . . 7~f7717 7/rr(7777r7777 ~ ~ ~ ~ ~ g0 ~ ~ ~ ~ ~ ~ ~ T I~~~ ~~ ~~
~ ~ ~ ~ ~ ~ 7777777 /7 I 77777 f7/ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ i I~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ /777771 7777177 m ~ ~ ~ '
~ ~ ~ ~ ~ I I ~' ~ ~ ~ ~ ~ ~ ~ 111"77 717777 ~ e ~ '
~ ~ ~ ~ ~ I I~ ~ ~ ~ ~ ~ ~ 7717 / 77771 9)999n9~ 777711 '
~ ~ ~ ~ ~ I I~ ~ ~ m ~ ~ 7/771 7777777771 9999 090999 7777/ 7 '
~ ~ ~ eef I ~ ~ ~ ~ ~ /77'17 1777777777'll 9999 99'/9 7'l7 7 ~ ~ ~ ~ ~ T I 177/1 177777777/7/7//7 ~99 9999 77 7/ ~ ~ e ~ I
] /777777771777/771717771 99 ', 9999 7 7f ~ ~ ~ I 1/ ~ ~ I 7/1771171771 '9~9 71 <> >>
I'SS.SS3 I
I 555
<<a<<a<<a<<a<<<<a<<a<<a<<a<<a<<a<<a<<a<<
5 1
57 a ~
9<99 9999
<<al <<a<<a<<a<<a<<aaaaaaaa
/7 I
HOT SIDE COLD SIDE Radial velocity (m/sec) at 91 in above tubesheet (CR = 3.76)
FIGURE B.26
APPENDIX C GINNA STEAM GENERATOR EXTERNAL TUBE LOADING TEST C. 1 ~Summar Work under this program is concerned with performing a detailed analysis of the effects of axial load and loose part impacting on the Ginna steam generator tube's propensity for local buckling.
Quantitative shot peening parameters such as shot size, velocity and duration will be developed which correctly simulate the tube distress caused by the loose metal piece found in the Ginna steam generator. Tubes representative of the various intensities of peening will be metallurgically examined to quantify the depth of affected metal and 'characterize the tube surface condition.
Examination will include light microscopy and microhardness of tube wall cross sections as well as scanning electron microscopy of the peened surface. In addition, destructive metallography will be carried out as part of this effort.
.2 ~ttd2 Two series of tests will be performed using the external tube loading test apparatus shown in Figure C.l. In the first test, the Axial Compressive Load Only Test, a tube sample with no peening is rigidly clamped at the tube sheet level with the first two supports positioned for a predetermined offset. An axial load is applied by mechanical torquing through a load cell until the tube sustains permanent deformation or until the maximum displacement due to differential temperature is reached. During the loading process, periodic data sets are taken. The test is repeated using offsets of zero, 1/4", 1/2" and other offset values as may be required.
The second test, Parameter Combination Test, involves samples that are peened at a predetermined amount. A predetermined axial load is then applied and the tube is subjected to external pressure until a sharp pressure reduction indicates local tube buckling or the maximum equipment test pressure capacity is reached. During the test, periodic data sets are taken at appropriate stop points in the loading sequence. Testing input variables will include:
a) Axial Load (1) Direction - tension or compression (2) Magnitude - from zero to 3,000 lb. using a load cell (3) Eccentricity - support offset, zero to 1/2 inch
b) Preen in (1) Area Covered (2) Shot size, shot velocity and shot duration based on comparison of RG&E input with peening bench test.
(3) Metal Removal - to be finalized later c) External Pressure External Pressure will be applied in increments up to a value exceeding the Ginna steam generator secondary side design pressure.
Measured output data is to include the following:
a) Forces developed at supports (leads to calculation of bending moment).
b) Tube bowing (deflections taken at various vertical positions).
c) Cross-section dimensions in peened region (micrometer and profilometer).
d) ECT readings to determine (1) loss of wall thickness, (2) section irregularity and (3) surface distress (background noise).
e) Load required to move "adjustable tube support" vertically 0 (tendency to lock up).
f) Visual and photographic record.
Tube collapse pressure (local buckling).
g)
C.1-2
0
~
LOATaaIG OCVICC (tVSII OR LOAD CELL ADAFfE1C TvsC WvaCO(teTR OACCT~
~
~ ALL SCARIIIG tlLLOV SLOCRS Set tORT MATC
~ 515 OO.TVSC StCCIVICN 1.15 OIA SIIAT'T 15 Leve t R C 5 5 VAI Z AT ION CIIAuSC R(AOVVSTASLQ
~ SWJ'T SVttORT SLOE'
~
~ ~
TvsC SaCCT CAISTIIIO I'SCA~ COIViVI IIOLOCR FLOOR EXTERNAL TUBE LOADING TEST APPARATUS COMBUSTION ENGINEERING FIGURE C.1