ML19308D660
| ML19308D660 | |
| Person / Time | |
|---|---|
| Site: | Crystal River  |
| Issue date: | 12/10/1976 |
| From: | FLORIDA POWER CORP. |
| To: | |
| Shared Package | |
| ML19308D659 | List: |
| References | |
| NUDOCS 8003120714 | |
| Download: ML19308D660 (240) | |
Text
- - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - - , - - - - - - - - - - -
O l
l 1
l
)
l 1
I i
6 i
(
8003120}{l L
EPOXYGRour LOWER , 'A 9-2/
DRIL L EPOXY v
, FROM GROUT HOL ES __ , , PL A CE REBAR MA TS v g 9-21
\
\ :
?
\ SET & GROUT RA DIAL R& BARS l v
POUR & CURE CONC. _ m sit _
si -
9-29 10 -1 11- 15 x
N D
- d "
! N N o S O <
o Q v v d
RertNsionius
. CR YS TA L RIVER N93 Nu cL E A R PL A N T IO *25(/8 - TENDONS) m DOME REPA/R ScuEcutC Ei 7l m
b
REACTOR BUILDING DOME DELAMINATION FLORIDA POWER CORPORATION CRYSTAL RIVER UNIT 3
. TABLE OF CONTENTS-Section Title Page TABLE OF CONTENTS i LIST OF TABLES iv LIST OF FIGURES v
1.0 INTRODUCTION
1-1 1.1 PURPOSE l-1 1.2
SUMMARY
AND CONCLUSIONS 1-1 2.0 ORIGINAL STRUCTURE 2-1
~
2.1 PHYSICAL DESCRIPTION 2-1 2.2 APPLICABLE CODES AND STANDARDS 2-2 2.3 CRITERIA 2-2 2.3.1 Flexural and Membrane Tensile Stresses 2-2 2.3.2 Shear 2-3 2.4
SUMMARY
OF ANALYTICAL RESULTS 2-4 2.4.1 Structure Prior to Operation 2-6 2.4.2 Normal Winter Operating Condition 2-6 2.4.3 Structural Integrity Test 2-6a h 2.4.4 2.4.5 Accident Condition Small Pipe Break 2-7 2-7 2.4.6 Summary 2-7 3.0 PROBLEM DEFINITION 3-1 3.1 INVESTIGATIONS 3-1 3.1.1 Original Indications _ 3-1 3.1.2 Preliminary Exploration 3-1 3.1.3 Additional Investigations 3-2 3.1.4 Other Events 3-2b 3.2 DELAMINATION GAP 3-3 3.3 POSSIBLE CAUSES OF THE DELAMINATION 3-3 3.3.1 Properties of Concrete and Constituents 3-4 3.3.2 Radial Tension Due to Prestress 3-5 3.3.3 Compression-Tension Interaction 3-6a l 3.3.4 Thermal Effects 3-7 3.3.5 "endon Alignment 3-8 3.3.6 Heavy Construction Loads 3-9 3.3.7 Coastal Location 3-9 3.3.8 Location Adjacent to Fossil Units 3-9 3.3.9 Construction Methods 3-9 3.3.10 Impact Loads 3-10 3.3.11 Shrinkage Effects 3-10
3.4 CONCLUSION
S 3-11 A.
1 Revised: 12-10-76
.. m a. .+au,, _. ..-~.--e.. s~-
g,
m TABLE OF CONTENTS (CONT'D)
Section ,
Title Page 4.0 DELAMINATED STRUCTURE 4-1
4.1 INTRODUCTION
4-1 4.2 APPLICABLE CODES AND STANDARDS 4-1 4.3 CRITERIA 4-1 4.4 EVALUATION 4-1 4.4.1 Structure Prior to Operation 4-4 4.4.2 Normal Winter Operating Condition 4-4a 4.4.3 Structural Integrity Test 4-5 4.4.4 Accident Condition 4-5 4.5
SUMMARY
AND CONCLUSION 4-7 5.0 CORRECTIVE ACTION 5-1
5.1 INTRODUCTION
5-1 5.2 REPAIR METHOD 5-1 5.2.1 Instrumentation 5-1 5.2.2 Dome Retensioning 5-3 5.2.3 Drilling Radial Holes 5-4 5.2.4 Delaminated Cap Removal 5-4 5.2.5 Inspection of the 24" Structure 5-4 5.2.6 Lower Level Crack Grouting 5-5 5.2.7 New Reinforcement 5-5 5.2.8 New Cap 5-5a l 5.2.9 Dome Retensioning 5-6 5.2.10 Structural Integrity Test (SIT) 5-6 5.2.11 Dome Surfacing 5-6 l
6.0 QUALITY ASSURANCE 6-1
7.0 REFERENCES
7-1 FIGURES APPENDICES Appendix A - Dome Concrete A-1 Appendix B - FPC Transmittal to NRC Dated July 11, 1974 B-1 Appendix C - Progress Report of Concrete Quality Evaluation C-1 Appendix D - Computer Program Verification and Description D-1 Appendix E - Photographs of Existing Structure E-1 11 Revised: 12-10-76
TABLE OF CONTENTS (CONT'D)
Section Title Pm APPENDICES (CONT'D)
. Appendix F - Discussion of Repair Method F-1 Appendix G - Comparison of Designs G-1
. Appendix H - Cadwelding Requirements H-1 Appendix I - Radial Tension and Shear I-l Appendix J - NRC Correspondence - SIT -
Instrumentation J-1 Appendix K - Discussion of Concrete and Aggregate Properties K-1 r-(
ii(a) Revised: 12-10-76
[- - - _. _ ,_.. . . _ . . . _ .____._.._.;. _ _
.- . .- . - - - . ~ . . - . - . . . . - . .1 TABLE OF CONTENTS (CONT'D)
Section Title Pye, SUPPLEMENT i - RESPONSES TO NRC COMMENTS TO THE CRYSTAL RIVER #3 REACTOR BUILDING DOME DELAMINATION REPORT S-1 GENERAL S-1 SECTILN 1.2 S-2 SECTION 2.3 AND TABLE 2-2 S-3 (Attachment i for answer to question 2.3(2) - 4 sheets)
SECTION 2.4 S-4 .
(Attachment 2 for answer to question 2.4(1) - 2 sheets)
(Attachment 3 for answer to question 2.4(4) - 3 sheets)
SECTION 3.1 S-6 SECTION 3.3 S-7 (Attachment for answer to question 3.3(3) - 1 sheet)
(Attachment for answer to question 3.3(6) - I cheet)
SECTION 4.4 S-9 SECTION 5.3 S-10 (Table 1 for answ ;r to question 5.3(6b) - 1 sheet)
SUPPLEMENT 2 - RESPONSES TO STRUCTURAL ENGINEERING BRANCH COMMENTS AND REQUEST FOR ADDITIONAL INFORMATION ON CR#3 DOME DELAMINATION INTERIM REPORT AND SUPPLEMENT NO. 1 1 I
GENERAL COMMENT
S 1 II DOME REPAIR 3 III CAUSES OF DELAMINATION 6 ATTACHMENT 1 J. T. Rodgers Letter (10-8-76) to Mr. John Stolz (1 sheet)
Instrument Location (2 sheets)
Displacement Acceptance Criteria (2 sheets)
J. C. King Letter (9-23-76) to F. L. Moreadith, Epoxy Grouting (1 sheet) iii Revised: 12-10-76
. _ _ _ . . . ~ . ._ ..
v*
x LIST OF TABLES
. Text Tables Page No. Cross References 2-1 DOME POUR LOG 2-8 2-1 2-2 DESIGN CRITERIA 2-9 2-2, 2-4, 2-6, 2-7, 4-1, 4-3, 4-4, 4-5, S-3, S-4, S2-6, S2-7 l 2-3 CONTROLLING LOAD COMBINATIONS 2-10 2-2, 2-3, 2-4, 2-5, 2-7, 4-1, 4-2, S-5 3-1 INITIAL EXPLORATORY CORE HOLES 3-12 3-1 -
3-2 CORE STRENGTH TEST RESULTS (3 sheets) 3-13 3-2, 3-4, 3-5, S2-5 l 3-3 CORE LOCATIONS AND MAIN DELAMINATION DETAILS (3 sheets) 3-16 3-2, 3-2a 3-3A RB DOME - SECONDARY CRACK LOCATIONS (2 sheeta) 3-18a 3-2a s
3B AVERAGE LIFT-OFF DATA 3-18e 3-2a, 5-6 3-4 DOME EP0XY AND CONCRETE STRENGTH TLSTS (3 sheets) 3-19 3-4, 3-6a l 3-5 DOME POUR LOG 3-22 3-9 3-6 PRESTRESSING SEQUENCES 3-23 3-9 3-7 ACTUAL TENDON STRESSING LOG 3-24 3-10, 3-18e l 3-8 TENDON STRESSING / CUMULATIVE DAILY TOTALS 3-25 3-10 5-1 DOME DETENSIONING SEQUENCE 5-7 5-3, S-6 ,
l l
l i
i w
l iv Revised: 12-10-76
LIST OF FIGURES Text Figure Title Cross References 2-1 REACTOR BUILDING DIMENSIONS 2-1 2-2* NON-PRESTRESSED REINFORCEMENT i 2-1, I-l 2-3. DOME CONCRETE PLACEMENT - PLAN 2-1, 2-8, 3-9 2-4 DOME CONCRETE PLACEMENT - SECTION 2-1 2-5 DETAIL "A" 2-1 2-6 5UPPORT DETAIL 2-1 2-7 DOME TENDON LOCATION PLAN 2-1
.2-8 CONDUIT SPLICES LOCATION & DETAIL 2-1 2-9** DOME DETAILS 2-1 2-10 36" DOME DEAD LOAD + PRESTRESS MERIDIONAL STRESSES 2-6, 3-6a l t 2-11 36" DOME DEAD LOAD + PRESTRESS HOOP STRESSES 2-6
_2-12 36" DOME DEAD LOAD + PRESTRESS MEMBRANE STRESSES 2-6, I-1 2-13 36" DOME DEAD LOAD + PRESTRESS LINER STRAIN 2-6 2-14 36" DOME DEAD LOAD + PRESTRESS SHEAR-FSAR 2-6, I '.
2-15 DELETED
. 2-16* 36" DOME NORMAL WINTER OPERATING MERIDIONAL STRESSES 2-6 2-17* 36" DOME NORMAL WINTER OPERATING HOOP STRESSES 2-6 2-18 36" DOME NORMAL WINTER OPERATING MEMBRANE STRESSES 2-6a 2-19 36" DOME NORMAL WINTER OPERATING LINER STRAIN 2-6a
! 2-20* 36" DOME NORMAL WINTER OPERATING SHEAR - FSAR 2-6a 2-21 DELETED 2-22* 36" DOME STRUCTURAL INTEGRITY TEST MERIDI'NAL O STRESSES 2-6a
- Revised: 8-10-76
- Revised: 9-22-76 4
v Revised: 12-10-76 sy,y 3 . _ _ _
- . . . . ~ ._ ~ -. _ ,_ ,-
t L
LIST OF FIGURES (CONT'D)
Text Figure Title Cross References 2-23 36" DOME STRUCTURAL INIEGRITY TEST HOOP STRESSES 2-6a 2-24 36" DOME STRUCTURAL INTEGRITY TEST MEMBRANE STRESSES 2 2-25 36" DOME STRUCTURAL INTEGRITY TEST LINER STRAIN 2-7 2-26 36" DOME STRUCTURAL INTEGRITY TEST SHEAR - FSAR 2-7 2-27 DELETED 2-28* 36" DOME 1.0 D + 1.0 F + 1.5 Pa + 1.0 T ,
MERIDIONAL STRESSES 2-7
.2-29* _36" DOME 1.0 D + 1.0 F + 1.5 P , + 1.0 T a HOOP STRESSES- 2-7 2-30* 36" DOME 1.0 D + 1.0 F + 1.5 Pa + 1.0 Ta MEMBRANE STRESSES 2-7
. g r n 2-31 36" DOME 1.0 D + 1.0 F + 1.5 Pa + 1.0 Ta LINER STRAINS 2-7 2-32* 36" DOME 1.0 D + 1.0 F + 1.5 P, + 1.0 T, SHEAR - FSAR 2-7
' 2-33 DELETED 2-34 36" DOME SMALL PIPE BREAK LINER STRAINS 2-7 3-1 INITIAL INVESTIGATION DETAILS 3-1 3-2 INITIAL INVESTIGATION CORE HOLE DETAILS 3-1 3-3 INITIAL INVESTIGATION - CORE HOLE DETAILS 3-1 3 LOCATION OF DEEP EXPLORATORY CORE HOLES 3-1 3-5 DEEP EXPLORATORY CORE HOLE DETAILS 3-1 3-6 DEEP EXPLORATORY CORE HOLE DEIAILS - 3-1 3-7 DEEP EXPLORATORY CORE HOLE DETAILS -
3-1 3-8 CONSTRUCTION JOINT CORE HOLE DETAILS 3-2
- Revised: 8-10-76 v1 Revised: 9-22-76 gge- n .~ --. -
1 LIST OF FIGURES (CONT'D)
Text Figure Title Cross References 3-9 CONSTRUCT " '9 INT CORE HOLE DETAILS 3-2 3-10 CONSTRUCTION JOINT CORE HOLE DETAILS 3-2 3-10A* FEATHER EDGE CORE PLAN 3-2a 3-10B* FEATHER EDGE CORE LOCATION IN " WINDOW" (TYPICAL) 3-2a 3-10C* SECTION - RB DOME (SHEET 1) 3-2a 3-10C* LOCATION OF MAJOR GAP (SHEET 2) 3-10D* LOWER LEVEL CRACK EXPLORATORY PROGRAM 3-2a, 3-18a 3-10E** SEQUENCES 1-6, DETENSIONED TENDON LOCATION 3-2a 3-10F**- 24" DOME DEFLECu0NS MEASURED AT LINER 3-2a 3-10G** LINER STRAIN AT 30'-0" RADIUS, MEASURED HOOP (R119H) 3-2a 3-10H** LINER STRAIN AT 30'-0" RADIUS, MEASURED MERID. (R119M) 3-2a 3-10I** LINER STRAIN AT 30'-0" RADIUS, MEASURED HOOP (R121H) 3-2a 3-10J**- LINER STRAIN AT 30'-0" RADIUS, MEASURED MERID. (R121M) 3-2a 1 3-10K** CONCRETE STRAIN AT TOP SURFACE OF 24" DOME, 30'-0" l RADIUS, MEASURED HOOP (152H) 3-2a I l
3-10L** CONCRETE STRAIN AT TOP SURFACE OF 24" DOME, 30'-0" I RADIUS, MEASURED MERID. (153M) 3-2a 3-11 DOME CORE HOLE PLAN NORTHWEST QUADRANT 3-2, 3-2a 3-12 DOME CORE HOLE PLAN SOUTHWEST QUADRANT 3-2, 3-2a 3-13 DOME CORE HOLE PLAN SOUTHEAST QUADRANT 3-2, 3-2a 3-14 DOME CORE HOLE PLAN NORTHEAST QUA11 RANT 3-2a .
3-15** RADIAL TENSION DUE TO PRESTRESS - ORIGINAL STRUCTURE 3-5, I-1
- Revised: 8-10-76
- Revised: 9-22-76
~
vii Revised: 9-22-76
. - a-
. \,-
LIST OF' FIGURES (CONT'D)
Text Figure Title Cross References
'3-16* AXISYMMETRIC FINITE ELEMENT MODEL 3-Sa, S-7 l 3-17*** STRESS CONCENTRATION IN TWO DIMENSIONAL STATE OF STRESS 3-6, l 3-18*** DELETED 4 3-19*** COMPRESSION - TENSION INTERACTION 3-6a 3-20 SOLAR HEATING TRANSIENT TEMPERATURE PROFILE 3-7 3-21 THER.M GRADIENT THROUGH THICKNESS AT CONSTRUCTION JOINT L-M DUE TO SOLAR RADIATION HEAT 3-7 I
- Revised: 8-10-76
- Revised: 12-10-76 I
l l
l vii(a) Revised: 12-10-76 ;
i
LIST OF FIGURES (CONT'D)
Text Finure Title Cross References 22* PLANE STRAIN FINITE ELEMENT NODEL 3-8, 3-11, S2-8 l 3-23 TEMPERATURE GRADIENTS DUE TO TENDON GRE. .J ,
3-8 3-24 DOME SECTION - 0 AZIMUTH 3-8, 3-10 3-25 DOME SECTION - 90 AZIMUTH 3-8, 3-10 3-26 DOME SECTION - 180 AZIMUTH 3-8, 3-10 3-27 DOME SECTION - 270 AZIMUTH 3-8, 3-10 3-27A*** TENDON D-118 (WEST SIDE) TOP OF CONDUIT PROFILE SHOWING RADIAL FORCE FROM TENDON MISALIGNMENT 3-8a 3-27B*** TENDON CONDUIT CURVATURE 3-8a 3-27C*** TENDON CONDUIT CURVATURE 3-8a 3-27D*** TENDON CONDUIT CURVATURE 3-8a 3-27E*** TENDON CONDUIT CURVATURE- 3-6a 3-28 SEQUENCES 1-6 SPECIFIED PRESTRESSING SEQUENCE 3-9 3-29 SEQUENCES 1-10 SPECIFIED PRESTRESSING SEQUENCE 3-9 3-30 SEQUENCES 11-21 SPECIFIED PRESTRESSING SEQUENCE 3-9
' 3-31 SEQUENCES 22-31 SPECIFIED PRESTRESSING SEQUENCE 3-9 3-32 SEQUENCES 32-41 SPECIFIED PRESTRESSING SEQUENCE 3-9 3-33 DOME TENDONS STRESSED AS OF 11-4-74 3-10 3-34 DOME TENDONS STRESSED AS OF 11-8-74 3-10 3-35 DOME TENDONS STRESSED AS OF 11-12-74 3-10 3-36 DOME TENDONS STRESSED AS OF 11-18-74 3-10 3-37 DOME TENDONS STRESSED AS OF 11-20-74 3-10 3-38 DOME TENDONS STRESSED AS OF 11-26-74 3-10 3-39 DOME TENDONS STREFSE' AS OF 12-2-74 3-10 m
- Revised: 8-10-76
- Revised: 12-10-76 viii Revised: 12-10-76 w, , , y ,. - _ _ . - . - -
LIST OF FIGURES (CONT'D)
Text Finure Title Cross References 3-40. DOME TENDONS STRESSED AS OF 12-3-74 3-10
, 3-41 DOME TENDONS STRESSED AS OF 12-4-74 3-10 3-42 DOME TENDONS STRESSED AS OF 12-6-74 (DAY SHIFT) 3-10
.2- t f* 3-43 DOME TENDONS STRESSED AS OF 12-6-74 .(NIGHT SHIFI) 3-10 l i
3-44 ACTUAL STRESSING SEQUENCE SUMMATION OF VERTICAL FORCES 3-10 !
i l
.- l l
i l
i I
l N.
viii(a) Revised: 12-10-76
~:w4em;w - - - . - - - . . . . - . . .. -.
LIST OF FIGURES (CONT'D)
Text Figure Title Cross References 3-45 ACTUAL STRESSING SEQUENCE SUMMATION OF VERTICAL FORCES 3-10 3-46* TANGENTIAL STRESS IN CONCRETE DUE TO SHRINKAGE 3-11 4-1* DOME MODEL 4-1, 4-2 4-2* 24" DOME DEAD LOAD + PRESTRESS MERIDIONAL STRESSES 4-4 4-3* 24" DOME DEAD LOAD + PRESTRESS HOOP STRESSES 4-4 4-4* 24" DOME DEAD LOAD + PRESTRESS MEMBRANE STRESSES 4-4 4-5 24" DOME DEAD LOAD + PRESTRESS LINER STRAIN 4-4 4-6* 24" DOME DEAD LOAD + PRESTRESS SHEAR - FSAR 4-4. I-2 4-7 DELETED 4-8* 24" DOME NORMAL WINTER OPERATING MERIDIONAL STRESSES 4-4a, 4-7
/
4-9* 24" DOME NORMAL WINTER OPERATING HOOP STRESSES 4-4a, 3-6a l 4-10* 24" DOME NORMAL WINTER OPERATING MEMBRANE STRESSES 4-4a 4-11 24" DOME NORMAL WINTER OPERATING LINER STRAIN 4-4a ,
4-12* 24" DOME NORMAL WINTER OPERATING SHEAR - FSAR 4-5, I-2 13 DELETED
. 4-14* 24" DOME STRUCTURAL INTEGRITY TEST MERIDIONAL STRESSES 4-5
' 4-15* 24" DOME STRUCTURAL INTEGRITY TEST HOOP STRESSES 4-5 4-16* 24" DOME STRUCTURAL INTEGRITY TEST MEMBRANE STRESSES 4-5 4-17 24" DOME STRUCTURAL INTEGRITY TEST LINER STRAIN 4-5 4-18* 24" DOME STRUCTURAL INTEGRITY TEST SHEAR - FSAR 4-5, I-2 4-19 DELETED 4-20* .?4" DOME D + F + 1.5P + T, (UNCRACKED ANAI.YSIS) 4-5, 5-5, Fig. 5-20 MEMBRANE STRESSES and 5-21 E,.
- Revised: 8-10-76 ix Revised: 12-10-76 mm, neem r . - --
( '
LIST OF FIGURES (CONT'D)
Text Figure Title Cross References 4-21 24" DOME D + F + 1.5P + Ta (UNCRACKED ANALYSIS)
LINER STRAIN 4-5 4-22 24" DOME D + F + 1.5P + T MEMBRANE STRESSES AT D + F + 1.15P - Ta(ASSUMfDCRACKING) 4-6 4-23 24" DOME D + F + 1.5P + T MERIDIONAL a LINER STRAINS 4-6 4-24 24" DOME D + F + 1.5Pa + Ta HOOP LINER STRAINS 4-6 4-25 24" DOME D + F + Pa+Ta MEMBRANE STRESSES 4-6 4-26 24" DOME D + F + Pa + T, (MAX. PREDICTED COMP. STRAINS)
LINER STRAIN 4-6 4-27* 24" DOME 1.0D + 1.0F + 1.5Pa + 1.01 , SHEAR - FSAR 4-6, I-2 4-28 DELETED 5-l** RADIAL HOLE LOCATION PLAN (NORTH HALF) 5-1, 5-4, 5-2** RADIAL HOLE LOCATION PLAN (SOUTH HALF) 5-1, 5-4 5-3* PLAN - NEW STRAIN GAGE GROUP LOCATION ON LINER 5-1, 5-3 5-4* CONCRETE STRAIN GAGE LOCATION PLAN 3-1, 5-3 5-5* CONCRETE STRAIN GAGE DCATION PLAN 5-1, 5-3 5-6* ' TlPICAL SECTION AT ILSTRUMENTATION LOCATION 5-1, 5-3 5-7* PLAN - EXISTING STRAIN GAGES ON OUTSIDE FACE REBAR IN DOME 5-1, 5-3 f
5-8* LINEAR POTENTI0 METERS FOR GAP MOVEMENT 5-1, 5-3 5-9** EXTENSOMETERS ON LINER FOR DISPLACEMENT MEASUREMENT 5-1, 5-3 5-10* TEMPERATURE MONITORING 5-1, 5-3 5-11** TYPICAL SECTION AT INSTRUMENTATION LOCATION (FOR SIT) 5-3
- Revised: 8-10-76
- Revised : 9-22-76 m
e x Revised: 9-22-76 nthsmx5:r; =~ ~ ~- ^~ ~' - --
LIST OF FIGURES (CONT'D)
Text Figure Title Cross References
. 5-12** TENDON LOCATION PLAN 5-4 5-13** TENDON LOCATION (ORIGINAL STRUCTURE) 5-4 5-14* TYPICAL SECTION - NEW CONCRETE 5-4, 5-5a 5-15** EDGE DETAIL 5-4, 5-Sa 5-16** GROUTING PACKER 5-5 5-17** PARTIAL PLAN 5-5 5-18** PARTIAL PLAN - SUPPLEMENTARY DOME REINFORCING 5-5, Fig. 5-19 5-19** SECTION AA 5-5, Fig. 5-18, S2-4 5-20** CAP REINFORCEMENT - MERIDIONAL DIRECTION PROVIDED
, VS REQUIRED 5-5, G-8, S2-4 5-21** CAP REINFORCEMENT - HOOP DIRECTION PROVIDED VS REQUIRED 5-5, G-8 5-22*** RADIAL STRESSES DUE TO PRESTRESS - DELAMINATED STRUCTURE 5-5a, S2-3, S2-6 5-23** RADIAL STRESSES DUE TO CREEP AND PRESTRESS -
REPAIRED STRUCTURE 5-Sa, S2-3 5-24** NEW CONCRETE POUR SEQUENCE 5-Sa 5-25*** DOME REPAIR SEQUENCE 5-1 5-26*** NO. 6 RADIAL REINFORCEMENT TEST 5-5
- Revised: 8-10-76
- Revised: 9-22-76
- Revised: 12-10-76 4
xi Revised: 12-10-76
~mmenn.u , m w. , - . --
j
1.0 INTRODUCTION
1.1 PURPOSE This report describes a delaminated condition of the dome of the Reactor Building of Crystal River Unit No. 3, the subsequent structural evaluation of the condition and the repair of the dome.
1.2
SUMMARY
AND CONCLUSIONS The condition was first discovered on April 14, 1976. Soundings, core borings and cutting investigations indicated that the dome had delaminated. The approximate maximum thickness of the delaminated concrete was found to be 15 inches with a maximum gap of approximately 2 inches between the two layers. The plan area of delaminated concrete was approximately circular in shape with a 105 foot diameter. The condition was not apparent via visual inspection of the dome surface. l Analyses of the delaminated structure were performed (see Section 4.0) and field investigations carried out (see Section 3.0) to determine its safety-and its capacity to sustain the design loads. Based on the analytical and field investigations, it was concluded that the structure in its delaminated condition did not represent a hazard to personnel. Furthermore, those investigations provided the basis for the repair method described in Section 5.0.
The repair of the dome was completed on October 30, 1976. The subcequent Structural Integrity Test (SIT) of the reactor containment building was successfully completed on November 3, 1976, the preliminary SIT report was filed with the U.CNRC Nov. 5, 1976 and the final SIT report on Dec. 7, 1976. The investigations and analyses presented herein and the SIT provide an adequate demonstration of the serviceability of the reactor containment building.
Potential contributing factors have been investigated in an attempt to determine the cause or causes of the delaminated condition. Several effects which may have contributed to the problem have been identified.
These include compression-tension interaction; tensile capacity of the concrete; misalignment of tendons; concentrated scresses generated by shrinkage, greasing and construction /. lays and stress concentrations associated with embedded c>nduit.
The calculated values of radial tension associated with several of the effects are as follows:
Radial Tension due to Prestress 41 psi (nominal)
(See Section 3.3.2)
Thermal Effects - (See Section 3.3.4) ,
- a. Sudden cool down 9 psi (nominal)
- b. Tendon' greasing 80 psi (peak) 1-1 Revised: 12-10-76 vrmwwwve ~- ~* ~
Tendon Alignment (Sce Section 3.3.5) 56 psi (nominal)
Shrinkage (See Section 3.3.11) 36 psi (peak)
In addition to the above, two local effects could have occurred at the construction joint between pours L and M due to construction delay (See
. Sections 3.3.4 and 3.3.11) . The radial tensions associated with the delay were predicted to be:
from Solar Radiation _ 280 psi (peak) from Shrinkage 360 psi (peak)
It has been concluded that radial tension stresses combined with biaxial compression to initiate the laminar cracking in a concrete hating lower than normal direct tensile capacity and limited crack arresting capability.
The procedures used in the investigations and results obtained are discussed in Section 3.0.
l l
l I- 1-2 Revised: 12-10-76 gr nruvrw .m,~rm hr= . ~ ~ .~ ~~~ '
c 2.0 ORIGINAL STRUCTURE 12.1 PHYSICAL DESCRIPTION The torispherical dome of the Reactor Building has a rajor inside radius of.110 ft - 0 in., a minor inside radius of 20 ft - 6 in. and a design concrete thickness of 3 ft - 0 in.. Lining 6e inside face of the dome is a continuous 3/8 in. thick carbon sted liner which acts as a vapor barrier. Figure 2-l'shows the basic configuration of the
, Reactor Building including the dome. The dome is prestressed by means of tendons forming a three way system. Non prestressed reinforcement i
as shown in Figure 2-2 is provided near the top surface in circumferential and radial directions. Also, shear and bottom reinforcement is provided adjacent to the supporting ring girder.
Concrete placement was symmetrical in the form of full depth concentric rings as shown in Figures 2-3 and 2-4, and Table 2-1.
The construction specification required that concrete be placed in a maximum of I ft - 6 in layers with the upper layers being vibrated into the lower layer to form a homogeneous full depth pour. An application of epoxy bonding compound was specified to be applied at each const:uction joint prior to placement of the next pour of concrete.
Support of the wet concrete was by means of ties from the dome tendon conduit to the angle anchors on the concrete side of the liner. The i e dome tendon conduits were tied together and carried the load back to previous concrete placements. See Figures 2-5 and 2-6 for details.
Concrete was specified to have a minimum 28 day compressive strength of 5000 psi.
Prestressing for the dome was 123 tendons arranged in a,three way (layer) system anchored at the ring girder. The 41 tendons in each layer were spaced at a horizontal distance of 2 ft 6 in., center to
. center. Refer to Figure 2-7. The conduit for each tendon was a 5 inch diametet schedule 40 galvanized pipe. Sections of conduit were i
joined together by means of a sleeve coupling welded to each section )
' ~
of conduit to form a grease tight system. The desired location of ,
spliced as shown on drawings and a typical splice detail are shown on 1- Figure 2-8. Each tendon has a guaranteed ultimate strength of 2.335 million pounds and was made of 163, 7mm diameter, low relaxation wires.
After the tendons were installed, the air space in the conduit uas ,
1 filled (bulk filled) with Visconorust 2090 P2 corrosion protection I grease. The grease was heated in a tank, and then pumped as a fluid
~
l
, via hoses, into the conduit. Based on information from the' '
constructor, the temperature of the grease at the tank outlet was i approximately 1600F.
The carbon steel liner plate, made from ASTM A 283 Grade C steel,
~
l lines the'inside surface of the dome. The liner also acted as the I inside form for the concrete. Attached to the concrete side of the '
l liner are steel angles which serve as anchors.
Figure 2-9 illustrates the basic dome details.
l i
2-1 A
.h $P fY' F 9 ^ 4 A& ,
[s 2.2 APPLICABLE CODES AND STANDARDS Codes and standards used in the design of the dome were given in the Crystal River Unit 3 FSAR (Docket No. 50-302), Chapter 5. The design predated the establishment of a concrete pressure vessel code. l Therefore the codes and standards used in the design and specified for the construction utre:
.a. Building Code Requirements for Reinforced Concrete, American Concrete Institute (AC':) 318-63.
- b. Specifications for Structural Concrete for Buildings, ACI 301-66 with modifications as r oted in the FSAR.
- c. Specification for the Lesign and Erection of Structural Steel for Buildings, 1963, AISC.
- d. ASME Boiler and Pressure Vessel Code,Section III, Nuclear Vessels,Section VIII, Unfired Pressure Vessels;Section IX, Welding Qualifications (applicable portions). -
- e. Specification for the Design and Construction of Reinforced Concrete Chimneys, ACI 505-54.
- f. AEC Publication TID-7024, " Nuclear Reactors and Earthquake," as amplified in the FSAR.
2.3 CRITERIA Table 2-2 lists the basic criteria applicable to the original design of the dome and Table 2-3 gives the controlling load combinations.
The design complied with the following' additional requirements as stated in the FSAR. .
2.3.1 Flexural and Membrane Tensile Stresses The allowable tensile capacity of concrete for membrane stresses (i.e., excluding all flexural and thermal stresses) due to the factored loads van 3/Td. The allowable tensile capacity of concrete for maximum fiber stresses due to the factored loads including the .
thermal load plus other secondary effects was 6/Tc . Where tensile fiber stresses exceeded the allowable, mild steel reinforcement was added on the basis of cracked section design. The amount of additional mild steel reinforcement and the increase in steel stresses due to temperature effects were determined in a manner similar to that I contained in ACI 505-54. The minimum steel on the exposed face of the concrete was 0.15 percent of the cross-sectional area of the concrete.
~
2-2 Revised: 8-10-76
- w .,wevmmv . . ~ ~ - , e-- - - - - - - - - -
r --
I s .
k-The concrete shell was prestressed sufficiently to eliminate tensile stresses due to membrane forces from design loads. Membrane tension due to factored loads was permitted to the limits described above. On those elements carrying primarily tensile membrane forces, any secondary tensile stresses due to bending could cause partial cracking. . Mild steel reinforcing was provided to control this cracking by limiting crack width, spacing, and depth. The capacity reduction factor "$" for tensile membrane stresses was taken as 0.95.
The coefficient "$" for flexure, shear, and compression is in accordance with ACI 318-63, Section 1504.
2.3.2 Shear In computing the shear capacity of the concrete the effects of membrane forces were accounted for as follows:
- a. When membrane tension exists or when membrane compression less than 100 psi, the section was designed to the ultimate shear provisions of Chapter 17 of ACI 318-63.
- b. When membrane compression of greater than 100 psi existed, the shear capacity was determined by the ultimate shear provision of Chapter 26 of ACI 318-63.
, The acceptability of the dome for shear is uvaluated by ,the following procedures:
The minimum capacity of the concrete, & Vc, is compared with the ultimate shear, Vu, that exists on the section. If a case results where this minimum concrete capacity is less than the ultimate shear, then the actual section capacity is computed using the FSAR criteria.
The shear capacities described above represent ultimate capacities.
. For assessing the section adequacy in shear, the ultimate shear Vu, is calcultted by applying a load factor of 1.5 to the net shears !
resulting from load combination a, b, and c in Table 2-3. A load factor of 1.0 is applied to the net shears .for load combination d.
I I
i l
l 4
I 1
s.
2-3 Revised: 8-10-76 s ,. e - ...--.m-m -- - - - - - - --- -- - - -- -
s ,
_(
When membrane compression greater than 100 psi exists, the shear l stress limits and shear reinforcing for radial shear used in the design were in accordance with Chapter 26, " Prestressed Concrete" of
-ACI 318-63, except as follows:
In equation (26-12) of ACI 318, the shear increment between l
flexural and diagonal tension cracking (0.6b'd /F )ewas modified based upon the results of testing under the direction of Professor A. H. Mattock of the University of Washington. The resulting equation is:
M Veg = Kgyb'd Q +M d+Vd v
~
3 where K3y = 1.75 0.036 + 4.0 np.
p i
i i
l l
v 2-3a Revised: 9-22-76
. ey,.pv i, .- ,=n ,
- ;-.- - - rn - - -v- - - - - -~ - - . - - -
h.
In accordance with ACI 318, the factor Kg is not considered to be greater than 0.6.
' Requirements for minimum shear reinforcement as called for in l Equation (26-11) of ACI 318 were provided only at discontinuities.
2.4
SUMMARY
OF ANALYTICAL RESULTS A summary of analytical results for the original structure is presented
's the FSAR. For purposes of comparison with the analytical results of the delaminated structure, results for the original structure are presented in this section. The format is consistent with that used in Section 4.4 for the delsminated structure. The original acceptance criteria used in the design is given in Table 2-2.
The controlling load combinations are given in Table 2-3.'
The structural analysis of the containment was performed using KALNIN'S Static Computer Program described in Appendix D. The individual loads which comprise the load combinations were input i
separately, and their results were combined internally in the program where possible. This was not possible for the Structural Integrity Test and Accident Condition load combinations due to the different Young's Modulus (E) values for the concrete under the sustained loads (D, .F, and T,) and the rapidly applied loads (P, and T ). In these 3
cases, stresses for each of the two types of loads were combined externally. The effects of shrinkage and creep were considered as discussed below.
- a. Shrinkage The effect of concrete shrinkage on the overall structural response (stress' resultants) is insignificant due to the large volume to surface ratios of the cylindrical wall, ring girder, and dome.
. In the prestress loss calculations, a conservative value for long term shrinkage strain of 100 micro in/in is used to be consistent with the original design. This is the value reconanended in Reference (12) for calculating prestress losses in fc = 5000 psi concrete. Actually, use of the shrinkage equation appearing in this reference for time = 40 yr. and volume to surface ratio =
24" results in a shrinkage strain of 10 micro in/in at end of plant life.
2-4 9-22-76 Re}tsed:
m, , , m-, w ,
-- J-- - -~ - --- - - - -- -
- b. Creep The effect of concrete creep under the prestress loads was
' included in the prestress loss calculations and in the structural analysis. The creep curves appearing in Reference (12) allow
- specific creep strains to be determined considering both concrete age at loading and duration of load. Actual creep strains were calculated from these specific creep strains for use in determining prestress losses. Also, the reduction in concrete stresses, which results in an increase in liner stresses, caused by concrete creep under sustained loads was taken into account in the structural analysis by using an effective Young's Modulus, EE. This modulus is expressed in terms of specific creep as E
E " 1 + sc Ee where:
Ee= instantaneous concrete Young's Modulus = 4 x 100psi, sc = specific creep (micro in/in/ psi)
~
Analysis of the containment for load combinations a, b, and c (Table 2-3) is based on calculated prestress losses and a sustained load (D, F, To) EE = 2.7 x 106 psi corresponding to the i present time. In load combination c (SIT), the results for 1.15Pa are based on Ee = 4.0 x 106 pai, l
1 For investigation of the containment under load combination d(LOCA),640 year calculated values of prestress losses and EE =
1.8 x 10 pai, are used. The 1.5Pa and Ta Part are based on Ec " l 4 x 106 pai, Prestress Losses
, c.
l The calculated prestress lesses (ksi) and effective prestress (ksi) are given below:
Elastic Steel Total Effective Shortening Creep Relaxation Shrinkage Losses Prestress Present Vertical 3.6 3.9 2.2 2.9 12.6 155.4 Hoop 6.4 7.0 2.2 2.9 18.5 146.25 Dome 6.9 7.6 2.2 2.9 19.6 148.4
- 1. 40 yr. l l s Vertical 3.6 9.1 3.4 2.9 19.0 149.0 Hoop 6.4 16.2 3.3 2.9 28.8 135.95 l Dome 6.9 17.6 3.4 2.9 30.8 137.2 2-5 mw w. . - -- - - - --
i
, ..s .
The membrane and extreme fiber stress results presented in this section are those obtained directly from the KALNIN'S Static Computer Program i analyses (i.e. linear, elastic, uncracked). At locations in the dome where tensile stresses exceed the allowable values ~given in Table 2-2, the concrete is assumed to be. cracked.. Cracked section investigations are performed to calculate concrete compressive and rebsr tensile
~
stresses. . In the cracking -investigation, the axial force (P) and moment (M) stress resultants applied on t'.te section are computed from the uncracked stresses (plotted). The only exception to this is for the Normal Winter Operating Condition load combination. In this case, the cracked section will reduce the effect of the through thickness gradient (AT) part of the T o term in the load combination. Therefore, the uncracked stresses due to AT are subtracted from the plotted stresses prior to computing P and M. Then, the effect of AT applied to the section with P and M is considered in a manner similar to that described in ACI 505-54. Cracked.section stresses, calculated as described above, are shown at selected locations in some of the !
figures for this section. Where allowable stresses are not exceeded.
existing compression reinforcement is not included in the cracking analysis.
- 2.4.1 Structure Prior to Operation For this load combination, the allowable extreme fiber stress according totheFSAR'requirementswas0.6fE=0.6x5000=3000 psi compression and 0 psi in tension. Results shown in Figures 2-10 and 2-11 give compression throughout the dome with a peak stress of 2666 psi in the neridional direction.
The maximum allowable membrane stress using the FSAR requirements was 0.45 fe = 0.45 x 5000 = 2250 psi compression and 0 psi tension.
Results shown in Figure 2-12 indicate a maximum membrane compression stress of 1836 psi. No membrane tension exists.
Uith reference to liner strains the limits are shown in Table 2-2.
Figure 2-13 shows a maximum compressive strain of 0.000948 in the meridional direction. For this load combination, the analysis indicates no liner tensile strains.
The shear strese limits noted in Section 2.3.2 were used in the
- original design. Figure 2-14 shows that the available shear capacity exceeds the ree.uired shear capacity using FSAR criteria.
2.4.2- Normal Winter Operating Condition 11e allowable extreme fiber stresses for this load combination
- . according.to the FSAR as given in Table 2-2 were 3000 psi in
-compression and 0 psi in' tension. F r a,n uncracked section, the results indicate that the stresses near the ring girder are tensile.
Thus, a cracked section investigation was required. This results in the peak comprescive stress of 3038 psi shown in Figure 2-16 and
_, 1022' psi shown in Figure 2-17. The rebar stresses indicated in these '
figures are considerably less than their 20 ksi allowable values.
2-6 Revised: 8-10-76
@M%5k:xk ' +"m ' ' ' -
'~*
, W ~~ ~ ~ ~ ~~' ~~ ~
l
('
Membrane stress limits as stated in the FSAR were 0.45 fE =
0.45 x 5000 = 2250 psi compression and 0 psi tension. The analytical l results shown in Figure 2-18 indicate a peak compressive stress of 1759 psi in the meridional direction, which is less than the allowable values. No membrane tensile stress exists. l Figure 2-19 shows the peak compressive strain to be 0.00124. For this-condition the analysis indicates no liner tensile strain.
For this load combination the shear capacity of the section is greater than the required capacity for the FSAR criteria. Refer to Figures 2-20.
2.4.3 Structural Integrity Test The extreme fiber stress allowable values noted in the FSAR are -
0.6 ff = 0.6 x 5000 = 3000 psi compression and 0 psi tension. The uncracked results shown in Figures 2-22 and 2-23 are a peak compressive stress of 1895 psi and a peak tensile stress of 227 psi.
Based on a cracked section investigation, a concrete compressive stress -
of 1930 psi exists in the outside face. The rebar stress on the inside face is not tensile, since the crack does not extend far enough into the section to reach the inside face rebar.
f
=
2-6a Revised: 8-10-76 J ,
_' . . _ ~
(
For the condition of membrane stress, the FSAR limit was 0.45ff=0.45x5000=2250 psi compression and 0 psi tension.
Figure 2-24 indicated the peak membrane compressive stress to be 834 psi, which is well below these values. No membrane tension exists.
Figure 2-25 shows a neak liner compressive strain of 0.00051.
Using FSAR criteria, Figure 2-26 shows that the shear capacity of the section exceeds the required capacity.
2.4.4 Accident Condition The allowable extreme fiber stresses according to the FSAR as given l in Table 2-2 were 3000 psi in compression and 6/Ej = 424 psi in tension.
Results shown in Figures 2-28 and 2-29 based on uncracked analysis indicates a maximum fiber compression stress of 1241 psi in the meridional direction. In accordance with the Design Criteria Table 2-2, rebar limited to a 36 kai design stress is provided where 6/El is e exceeded.
According to the FSAR requirementa, the maximum allowable membrane stresswas0.45ff=0.45x5000=2250 psi compression and 3/le =
212 psi tension. Results shown in Figure 2-30 inoicate a maximum concrete tensile stress of 148 psi. Therefore, no rebar was required to resist these membrane stresses.
- Figure 2-31 shows a maximum compressive strain of .002057. For this load combination the analysis indicates no liner tensile strains.
Using FSAR criteria, Figure 2-32 shows that the available shear capacity exceeds the required capacity.
2.4,5 Small Pipe Break Although not addressed in the FSAR, NRC personnel informally requested that the maximum liner strain be calculated for a small pipe break producing a liner tamperature (T,) of 3200F, This load combination is defined in Table 2-3. Analysis of this accident situation assumed to occur at the end of plant life indicates in Figure 2-34 a maximun
. compressive liner strain of 0.0029.
2.4.6 Summary The original design met the requirements set forth in the FSAR. It should be noted that the 90 day concrete strength equaled approximately 6000 psi thus resulting in further margin.
9
_\.*_
2-7 Revised: 8-10-76
-* Q*fitpMw .?O 7", - ry + - + +- re ~~M ' e.=-4.~ ~- .-----.~~-~,--v - .-- - - -
/
- TABLE 2-1 DOME POUR LOG Pour No. Date Location
- G1 2-18-74 45 - 135 G1 2-20-74 225 - 315*
G2 2-25-74 135 - 225 G2 2-25-74 315 - 45 H3 3-1-74 180 - 270 H3 3-4-74 0 - 90 H4 3-6-74 90 - 180 H4 3-7-74 270 - 360 JS 3-12-74 45 - 225
. J6 3-14-74 225 - 45 K7 3-20-74 135 - 315 K8 3-26-74 315 - 135 L9 4-4-74 0 - 360 M10 7-8-74 0 - 360 N11 7-12-74 0 - 360 P12 7-17-74 0 - 360
. Q13 7-22-74 0 - 360
- Refer to Figure 2-3 for plan location.
u' 2-8
@!.krgny u.v.:,; O .
i TABLE 2-2 DESIGN CRITERIA Original Delaminated Criteria Dome Dome Design accident pressure 55 psig 55 psig Design accidaat temperature 281 F 281 F Concrete compressive strength (ff) 5000 psi 6000 psi Allowable membrane compression stress 0.45fd = 2250 psi 0.45fl = 2700 psi Allowable extreme fiber compression stress 0.6fc = 3000 psi 0.6f4 = 3600 psi Service Loads
- Allowable membrane tension stress 0 0 Allowable extreme fiber tension stress 0 0 Factored Loads * ,
Allowable membrane tension stress 3/f'c = 212 psi 0**
Allowable extreme fiber tension stress 6/fc = 424 psi 0**
Liner ***
Compression strain .005 in/in ASME Section III Tension stress fy Division 2
- Mild steel reinforcement (Grade 40) is provided where the allowable stresses are exceeded. The stress in this reir's ;ement is not to exceed 20 ksi for Service Loads and 36 ksi for Factored Le
- The concrete tensile capacity it enservatively assumed to be zero for factored loads in the delaminated dome e aation.
- The design utilized 0.84 f for y both compression and tensile stresses during construction.
L-2-9 Revised: 8-10-76 ww.mm esn- m~ ?:-v ~r~~? ,- m-~~-~- - - ~ ~ ~ ~ - -
TABLE 2-3 CONTROLLING LOAD COMBINATIONS
- a. Structure Prior to Operation 1.0D + 1.0F
- b. Normal Winter Operating Condition 1.0D + 1.0F + 1.0To
- c. Structural Integrity Test 1.0D + 1.0F + 1.15Pa + 1.0Tt
- d. Accident Condition 1.0D + 1.0F + 1.5P, + 1.0T a
- e. Small Pipe Break Accident 1.0D + 1.0F + 1.0T a Symbols used in the equations are:
D = dead load of structure F = prestress force To = operating temperature Te = test temperature Pa = accident pressure T, = accident' temperature including To L.
2-10 Revised: 9-22-76
.cw y + a m & vn y - ~r - -
e
. 3.0' PROBLEN DEFINITION 3.1- INVESTIGATIONS 3.1.1 Original Indications l
On April 14, 1976 electricians were attempting to secura drilled-in anchors to the top surface of the done and certain anchors would not hold. on the west side of the done, 52 ft - 6 in. from the apex, ,
further investigation revealed an area of the dome surface which sounded hollow when hit with a hanumer. The size of the area was approximately 30 feet long by 3 feet wide. The area was curved slightly
, as it appeared to follow the circumferential construction joint between i t pours G and H. This area is shown in Figure 3-1.
In order to better define the hollow area concrete soundings were employed. The results of this examination indicated that a hollow sounding zone encircled the done. No cracks appeared on the done ,
- surface. The plan of the outer bounda y of the affected area as '
determined by concrete soundings is shot , in Figure 3-1, a
3.1.2 Preliminary Exploration i
To determine the extent of the condition, it was decided to examine a portion of the initially discovered area. The examination consisted j
,-. of saw cutting a 2 ft-6 in. long by 10 in, wide by 3-1/2 in. (maximum) '
deep section of concrete. Inspection of the removed pieces of ;
concrete as well as the lower surface revealed a clean fracture (100% i fracture of coarse aggregate) with little or no powder or rubble in the joint. The exposed surface revealed a gap running up into the dome I towards the apex. A three foot length of wire could be probed .into the gap witheut resistance. This gap, coupled with the results from the concrete scundings, appeared to indicate that the rop of the dome had l delaninated. The saw cut examination area is shown in Figure 3-1. '
To confirm that the dome was delsminated, seven exploratory core holes i were drilled. The location of each core hole was carefully selected to ensure that the stressed tendons were not damaged while achieving sufficient coverage of the dome to establish the extent of the
, condition. The location of these core holes is shown in Figure 3-1, with details shown in Figures 3-2 and 3-3. The results of the i exploration, sumanarized in Table 3-1, shows that the dome had
- delaminated. The thickness of the delsminated concrete was found to be approximately 15 in. in the area of the apex. The dimension of the
, gap in the same area was approximately 1-3/4 inches. See Section 3.2 fot a discussion of calculated gap versus measured gap.
~
The seven exploratory core holes were shallow holes confined to the i thickness of the delsminated cap. By maintaining a shallow depth,
- contact with tendon conduit was avoided. To determine if there was
' deeper cracking, eight core holes were drilled to an average depth of 29-1/2 inches. The location of these core holes is shown in Figure 3-4, with details shown in Figures-3-5, 3-6 and 3-7. Care was j 4
- taken to avoid the tendon conduit and to not penetrate the liner. Two 1
l l
3-1 Revised: 8-10 d. *
((sg% & .55e. : giff!?W pQi%-n.**T'?W W Y'L * ' ~ ' ""'W~~' "W ~*~: ~'
~'"":~"" '~
v+ w y , .c , # . -- y . - . ,w, -. ~.m. , y,,, - - ,,..,-e. , - - , , - , , _ - , , , . - , -,+--,,e,,w,
diameters of core hole were used, the first 12 to 14 inches of each core being 4 inches in diameter, and the remainder 1-3/4 inches in diameter. A borescope was then used to examine the surface of the
.1-3/4 in diameter hole. In five of the eight holes, additional cracks were found. The maximum width of thase cracks at the exposed surface ranged from approximately 0.01 inches to 0.03 inches, with the orientation of the cracks being parallel to the surface of the dome.
During core drilling, the drill water did not run out, indicating that the cracks might not be continuous. l With the knowledge that the dome had delaminated, further investigation was directed to determine the safety of the dome. A dimensional monitoring program measuring gap depth, dome cap displacement and ambient temperature was established. As results became available, they indicated thet the gap increased and decreased in delayed response to ambient temperature changes. Results of the monir. ring program, together with analytical re~ults s for the de1.4minated condition indicated that the structure was safe.
To provide the basis for determining the in-place strength of the dome "
con; tete, cores were taken from the dome cap. For compressive testing, two cores per pour were tested which represented one set per 100 cubic yards or lesser quantity placed per day. In addition, a third core per pour was taken and either saved or used for a split tensile test. This program required a total of 51 cores.
Table 3-2 lists these cores and their strength test results and Table 3-3 includes the location of these cores.
3.1.3 Additional Investigations In order to further investigate the c'ondition of the existing structure, the following programs were undertaken:
- a. Construction joint adequacy.
- b. Location of tendons.
- c. Direct tensile tests of concrete.
- d. Location and definition of the perimeter of the main delaminacion.
- e. Location and details of lower level cracks.
l
- f. Tendon forces, i.e., lift-off tests.
- g. 15% detensioning strains and deformations.
l To establish that construction joints were sound, a series of cores were made and examined. These cores (1, VIA, IIIA, IIIB, VB and IV in Figures 3-11 and 3-13) indicate that the epoxy joint material had effectively bonded the joint together and that an i . . acceptable bearing condition exists through the joints. Figures 3-8, 9 and 3-10 indicate details of these cores.
3-2 Revised: & .1.0-76 4 f .r-w 4 k.j. ..
In order to provide a definition of the existing structure in the vicinity of the perimeter of the main delamination, a series of core holes were made in pours J. H and G to determine the elevation of the main delsmination as well as the existence of any lower level cracks in pours (G and H) which contain radial reinforcing. Figures 3-10A and 3-10B indicate the location of these cores. Deep cores were also made l in the region of the construction joint between pours H and J to locate any lower cracks or concrete crushing that might have existed in this high compressive stress area. A description of the cores is contained in Figure 3-10C.
The cores indicate that the main delamination did not penetrate the reinforced areas of pour G and 9 and that lower level cracking also did not occur there. The observed condition of the concrete material in the deeper cores did not indicate any crushing. l An additional set of deep cores was drilled in each of the original pours to supplement the original data on deep cracks. Figure 3-10D and Table 3-3A indicate the location of the cores and cracks. The plan etrent of the cracks appeared to be rather extensive, at a relatively l tew levels. The levels appeared to correspond with the location of the tendon groups.
A sample group of 21 tendons were selected for lift-off readings in l order to determine the level of prestress on the dome and to establish s the accuracy of the predicted losses.
In addition to these lift-off tests, 18 other tendons (6 in each group) were detensioned and lift-off readings obtained. See Figure 3-10E. The detensioning of these tendons represents a symmetrical and uniform removal of 14.6% of the dome prestress. The associated stresses and deformations for this loading were compared with predicted values. Figures 3-10F thru 3-10L show these values. These figures indicate tLat the structure responded as a full 24 in deep structure -
since actual response was less than predicted response. The 14.6%
, removal of dome prestress was equivalent to applying an internal pressure of approximately 14 psig.
Table 3-3B indicates the average lift-off values and the predicted ,
design value for the 39 tendons. These results indicated that the I inplace domo prestress was 9% in excess of the predicted.
Details of the location of all cores in the general gap study are given in Table 3-3 and in Figures 3-11, 3-12, 3-13 and 3-14.
At no time during the investigation was there any evidence of grease l leakage from the tendon conduit.
l t - (
3-2a Revised: 12-10-76
-+ mere wmmn,- - - -,-- -~ ~ - -- -~
l
3.1.4 Other Events Exwination of construction records revealed that at 7:20 a.m. on December 4, 1974 a loud noise or boom occurred on the site. The noise reportedly appeared to come from the Reactor Building and was heard by certain construction workers in their construction change areas. One e
worker was in the Reactor Building personnel lock at the time, felt vibrations and saw dust falling. Since subsequent visual inspection by construction personnel of the Reactor Building interior and exterior including the dome did not reveal any damage, the incident was not reported to the NRC or the Engineer.
o
=
3-2b Revised: 8-10-76 KPMAM ~hd# drx*.:1.;$$'s)'. 4*sd Lj,y a~
~
,*"-T.Tr '~~" ~
"*?'"'"
~ ' ' ~ ~ ~ * * ~ ~ " ~ ~ ' ~ ~ ' ~ " ^ ~ ~ ~ ~ ~ ^ ~ ~ ^ '~~~~
(
3.2 DELAMINATION GAP Calculations were made to determine the apex displacement of the delsminated dome. The purpose of the calculations was to determine if correlation with the actual condition could be obtained. After allowance for concrete age of approximately 700' days and creep of the concrete, the delamination gap was calculated to be approximately 1-1/2 inches.- The actual gap averages 1-3/4 inches. The calculated displacement is for a 2 foot thick dome. The correlation between actual and theoretical gap depths was considered satisfactory.
3.3 POSSIBLE CAUSES OF THE DELAMINATION The engineering investigations took into consideration all factors which were believed to be potentially contributory to the delaminated condition. The source of much of the information was the constructors records. Factors considered in the investigation were:
- 1. Properties of concrete and constituents.
- 2. Radial tension due to prestress.
- 3. Compression - tension interaction.
- 4. Thermal effects.
(
- 5. Tendon alignment.
- 6. Heavy construction loads.
- 7. Coastal location.
- 8. Location adjacent to fossil units.
- 9. Construction methods.
- 10. Impact loads.
- 11. Shrinkage effects.
The following sections present discussions of these factors.
E v
3-3 Revised: 12-10-76
\
- -- g-mnemww,w, w --- ---~~~~- - - - - - - - - - - -
3.3.1 f roperties c ? - Concrete a- Constituento Sources of information on the properties of the concrete and its.
-constituents includa : results of tests performed for'the original mix qualification 3 supplemental tests, during construction and subsequent to discovery of the delaninated condition. l
- a. Oriainal Tests At the beginning of the project, special tests were conducted on a mix to determine'such properties as creep, shrinkage and thermal conductivity. In addition, concrete qrlinders were tested in compression and splitting tension. The average compressive strength of two sets of cylinders (three per set) tested at 28 and 56 days was 7040 psi and 7610 psi, respectively. Split tensile tests for two sets of cylinders (three per set) gave an average of 585 psi and 600 psi when tested at 28 and 90 days, respectively.
Test results are contained in Appendix A. l Requirements for concrete materials and available qualification test results are also contained or referenced in Appendix A.
The mix proportions for the done concrete are described in Appendix A.
( b. Construction Testing l
. During concrete placement of the dome, test cylinders were made.
The compressive strength of the cylinders was assessed in i accordance with the requirements of Section 4.3.3 of ACI 318-71 l and resulted in a 90 day design strength of 5930 psi. Table 3-4 l summarizes the results for both concrete and the epoxy bonding
! compound, i
On April 4, 1974 concrete was placed in pour L-9 of the reactor building dome. The concrete was to have a design strength of 5000 psi at 28 days. Seven day test cylinder breaks indicated that the concrete would not achieve the design strength level.
The 28 day cylinde7 ten results averaged 4570 psi. A review of the dome design was made to determine a minimum acceptable concrete strength for this particular area of the structure. The results of the review indicated that 4500 psi was an acceptable
, minimum design strength. Refer to Appendix B for further details.
1 The 90 day cylinder test results averaged 5820 psi.
c.- Subsequent Tests Subsequent to the discovery of the delamination, cores with a 3-3/4 in. diameter were obtained to establish in-place
-compressive strength. Locations of the cores are deceribed in
+
Section 3.1. The compressive strength of the cores was assessed s_c in accordance with the requirements of Section 4.3.3 of ACI 318-71 and resulted in a design strength of 6130 psi.
Results a: summarized in Table 3-2. Cores were also obtained i 3-4 Revised: 12-10-76
+ -;mmpeenmrecru~~r-wry-~r--m -
5 .m -
-- --- 1
k-. for testing by the split cylinder method. Locations of the cores are described in Section 3.1. Testing by this method indicated an average split tensile strength of 708 psi. Results are summarized in Table 3-2. The age of the concrete at the time the cores were tested ranged from approximately 630 days to 790 days with a weighted average age of 720 days.
Cores were also obtained for testing in direct tension to establish correlation between the standard tests, i.e.,
compressive end split tensile strength, and the direct tensile strength. _The average direct tensile strength of 420 psi when compared to the average compressive strength of 6177 psi for the poura from which tensile specimens were obtained and an average split tensile strength of 708 psi is low using standard ratios.
The ratio of the split tensile strength to the compressive ,
strength does not appear abnormal. When the range of tensile l values is considered, 230 to 505 psi, depending on the amount of soft aggregate, the variation from the normal ratio of strengths is pronounced. l 1
Based on the average compressive cylinder tests a design f' of 6000 psi is justified.
Concrete specimens were also petrographically examined.
(~ Results of the foregoing tests are presented in Appendix C.
3.3.2 Radial Tension Due to Prestress Any change in tendon direction produces radial stresses when the tendon is tensioned. Since the tendons are not at the outside surface of the dome the radial forces are tensile for the cone. 2te above the tendon groups and compressive for concrete below the tendon groups. ;
Subsections a, b and c address design criteria, material properties and {
stress concentrations in relation to radial tension due to prestress. j l
. a. Design Criteria j l
At the time of the structural design of the CR3 containment, there were no code criteria for allowable radial tension stresses. l Based on calculations, the average radial tensile stress at the centerline of the top tendon group (i.e., at the level of the delamination) for the theoretical geometry is approximately 41 psi based on the gross area (see Figure 3-15). The minimum direct tensile strength of the dome concrete, based on testing of cores (see pg. C-15), was 230 psi.
Delamination of the dome would not have been expected based on this comparison.
u.
3-5 Revised: 12-10-76
- rr.wew w. esswnew mr --
( b. Variation of Material Properties In order to evaluate the effects of material properties on radial tension stresses, the axisymmetric element in the SAP IV computer program was used to perform a parametric study. The basic model as shown in Figure 3-16 divided the through thickness shell into layers with the following properties. The concrete above and below the tendon groups was assumed to have a Young's modulus E =
4 x 106 psi and a Poisson's ratio V = 0.2. For the middle layer containing concrete and tendon conduit the assumptions were E =
3.17 x 106 psi and v = 0.2. The Young's modulus, E =
3.17 x 106 psi, was estimated on the basis of the ratio of net concrete volume to gross concrete volume. Values of E = 30 x 106 psi and v = 0.3 were assumed for the steel liner. In order to simulate the effect of rebar near the top surface, the equivalent Young's Modulus based on transformed area was used for the concrete near the top surface. To simulate the effect of conduit on Poisson's ratio, V = 0.3 was used for the middle layer. A lower bound value of E = 2.5 x 106 psi was also used for the middle layer to check the sensitivity of the structure to an assumed soft layer. The Poisson's ratio of concrete varies 3-Sa Revised: 12-10-76 c"M. %.. .
-. , .r P *-~2 "*"' *'7w w'a -M-> ; ~V** M.w=mn"*~~""~ '~
between 0.15 and 0.2 for stress below 40% of ultimate strength (3) .
Another reference gives test results which indicate an average value of v = 0.2(4). In order to simulate the effect of a concrete stress higher _than 0.4 f' v = 0.3 was assumed.
A summary of the parametric study is shown in the following Table:
'f Temperature Poisson's Young's Poisson's Rebar Ratio Modulus Ratio Case Basic Case Case Effect Effect Effect i
6 6 6 6 6 E (concrete) 4 x 10 4 x 10 4 x 10 4 x 10 4 x 10 E (concrete & 6 conduit) 3.17 x 10 6 3.17 x 10 0 3.17 x 10 0 2.5 x 10 3.17 x 10 6 v (concrete &
conduit) 0.2 0.2 0.3 0.2 0.3 v (concrete) 0.2 0.2 0.2 0.2 0.3 Rebar Effect No Yes No No No Resulting Max.
Radial tension 42.5 43.5 41.0 44.6 42.8 These calculations indicate that the radial tension stress is low and not sensitive to variations in these parameters.
2
- c. Stress Concentrations The geometry of the tendons in the CR3 dome suggests that stress concentrations around the conduit might be a centributing factor to the delsmination. Consequently, investigation of the potential for stress concentrations was undertaken.
Referring to Figure 3-17, a model is shown which represents a simplistic simulation of the conduit effect. If the hole were unlined (i.e., no Schedule 40 pipe), a theory of elasticity solution for the model shown, subject to a uniform uniaxial compression, oc , would result in a transverse tension at the face
( of the hole of ce=c. e That is, an applied compression of 1000 psi would result in transverse tension of 1000 psi.
In addition to the above, a detailed finite element analysis of the model subject to c with e a Schedule 40 pipe embedded as shown a was also investigated. Assuming linear elastic behavior and that the pipe and concrete remained bonded together, the results were somewhat different. The location and magnitude of maximum tensile stress was changed. The location and orientation of at is as shown la Figure 3-17. The magnitude of at was found to be approximately 0.5 o c. .
q 3-6 Revised: 12-10-76
-w - , m% m,~.--m. # -~m. :- , m m _ . _ - , . - - - --
-. _ , _ . _ . . . - . . _ _ _ . . ~ . _ _ - . . . . _ , , - _
a i
Available literature (e.g. References 13 and.14) and test results presented in Reference 15 indicate elastic stress concentration
. calculations do not accurately predict the failure stress for concrete structures. The effect of stress concentrations in the CR3 dome, however, may have been greater than would have been expected in normal concrete as a result of aggregate quality (see Appendix K).
2 3.3.3 Compression-Tension Interaction
.The state of stress at the top tendon layer is triaxial, two membrane
-compressive stresses and one radial tensile stress. It is conservative to.considet the interaction of one compressive stress with the radial tensile stress (Ref. 4, 8, 9, & 10). Before discussing compression-tension interaction in the CR3 dome at the time of stressing the tendons, it is necessary to establish applicable stresses and material strengths.
The initial membrane compression stre9ses would have been higher than those defined in Section 2.0, since the time dependent losses assumed in the calculations of Section 1.0 would not have occurred. The maximum meridional compressive stress for a fully prestressed dome (Figure 2-10) of 2,275 psi on the outside face would have been approximately 2,500 psi due to initial prestress forces. The radial tension stress could have been as high as 55 psi.
x '~
At the tile of stressing the tendons, concrete compressive strength
- (f') was approximately 6,000 psi (based on 90-day strengths, see l TaEle3-4). The splitting tensile strength of the material was in the neighborhood of 500 to 600 psi (based on trial mix data, see page A-3-2). The direct tensile strength based on recent tests (see page C-15) could have been as low as 230 psi for some of the dome l concrete. I I
Using the interaction lines suggested in Reference 15 (Figure 4-9), j the straight line failure envelope shown in Figure 3-19 was constructed I as follows:
1
- 1. A tensile strength of 160 psi was calculated using the minimum CR3 l tensile-test result of 230 psi with an approximate 30% reduction due to the presence of holes (see Reference 15, page 4-10).
2.- Acompressionstrengthof4800psiwascalculatedusingff=
6000 psi and a 20% reduction for size effects (see Reference 15, page 4-10).
~
The compressive strength was not reduced for the effect of holes since the transformed area of the pipe effectively fills the. holes for compressive stresses.
The_ interaction line used is more conservative than other possible representations of the failure envelope.
s.
i' 3'-6a Revised: 12-10-76 i
- 1 w9kQ pV9f* q .
7 - , , --,wr,.N N FEWW.**W VfCM'"W# 1 " " '~
_ - . . . _ _ . . _. - _ . _.- ,_- . . _ _ _ . . -,_,,--.,..m . - - - . , . . _ . . . . _ . _ , .
e .
The plotted state of stress for the dome (55 psi, 2500 psi) corresponding to full initial prestress indicates a nominal safety factor against delamination. However, the 55 psi radial tension stress due to prestress could have been increased in local regions of the dome due to o:her effects discussed in Section 3.3 and could have resulted in dalamination.
A compression-tension :.nteraction of the stresses in the concrete would explain the appearance of the main delamination surfaces, the delamination of the upper cap snd the presence of some cracking in levels below the top tendon group. It would also explain the fact that the structure retained its load carrying capacity subsequent to the delsmination.
s.
3-6b Revised: 12-10-76 mv , www , s-= n ~~n .
.n v,~- n ~ . m n- - - - --- -
- 3. 3.4 - Thermal Effacts Two types of thermal effects considered were solar radiation (environmental) and tendon greasing (bulk filling).
- a. Solar Radiation The effect of solar radiation on the surface concrete temperature was calculated. In perforr.ing this calculation the initial condition assumed was 60 F throughout the dome thickness. The effect of solar radiation was calculated to heat the dome surface to 1520F. Subsequent to a six hour heat up period, the 6.0 hour0 days <br />0 hours <br />0 weeks <br />0 months <br /> gradient shown in Figure 3-20 was calculated. To determine if a thermal shock could have had a significant effect on the stress state in the dome, a sudden cool down due to a thunderstorm was postulated. Therefore, a step function of a six hour quench using a surface temperature of 50 F0 was assumed.
Figure 3-20 shows the gradients after the initial heat up (0.5 hr),
just prior te quench (6.0 hr), after quenching (6.5 hr), and two points along the cooling period (8.0 hr and 12.0 hr).
Using the analytical model described in Section 3.3.2, and the gradients shown'in Figure 3-20, the maximum tensile stress at the level of the centerline of the top tendon group was calculated to be 8 psi.
The solar radiational heat also had an affect at construction joint L-M during the three month construction delay. The conduit i protruding from the joint had a different temperature than the l surrounding concrete. This causes hoop tension around the I conduit in the same way as that due to hot grease injection. The temperature gradient is as shown in Figure 3-21 and the maximum tension as calculated by plane strain element of computer program SAP IV (see Appendix D) is 280 psi.
Based upon these studies it is unlikely that the solar effect by itself could have produced other than very limited cracking at the construction joint interface.
- b. Tendon Greasing The field records show that tendon gressing took place in two stages. Eight unstressed tendons were greased prior to stressing the tendons (with the exception of three). The remainder were I -
- k. .
3-7 Revised: 12-10-76 l-mm, peri % m- - m--~mewwro. .- . -
. r- v ~ ~~- -
, greased about four months later, after all tendons had been stressed. .The second greasing operation was completed in a period of eight days.
The grease was heated prior to injection to reduce its viscosity.
According to available field records, the temperature of the grease at the tank outlet was in the range of 150-1700F, It was then pumped via a rubber hose into one tendon conduit it a time.
After all the air had been purged from the conduit, pumping ceased and the conduit was sealed at 0 psig.
During hot grease injection, the conduit heated more rapidly than the concrete due to the lower thermal conductivity of concrete.
For a step change of temperature from conduit to concrete, the tensile stress in the concrete surrounding the conduit can be calculated from the theory of elasticity. Using the compatibility of radial displacements at the conduit-concrete interface,(ll) the tensile stress is determined to be 11 psi /0F. The plane strain element of the SAP IV program as shown in Figure 3-22 yields a tensile stress of 15 psi /0F for the identical condition.
Based upon an averaging of field records of grease temperature in the storage tank and at the conduit outlet, a heat transfer analysis was performed to establish the thmperature gradients within the structure due to the greasing operation. The r resulting gradients at a point approximately midway between ring
- girder and dome apex are shown in Figure 3-23. Using the plane strain element in the SAP IV computer program, shown in Figure 3-22, the maximum tension stress is approximately 80 psi, which occurs at the location where the thermal gradient drops to zero.
It is recognized that variations in the greasing operation could produce more severe gradients and consequently higher concrete tensile stresses.
1 The injection pressure of hot grease can also cause tension j around the conduit. With the same kind of radial displacement compatibility calculation, the maximum grease pressurc reported ]
, 1 by the constructor of 85 psi causes 40 psi tension in the l concrete by compatibility calculation (ll). I The cases studied indicate stresses of sufficient magnitude to cause the delamination only when considered in conjunction with other effects.
3.3.5 Tendon Alignment (-
To establish the accuracy of the positioning of the tendons within the dome, a survey was conducted to determine the actual dome thickness and the depth from the dome exterior surface to the top of the upper tendon group conduit. The results of the survey are shown in Figures 3-24~through 3-27. This shows that the conduit are high near the l periphery and low at the apex and suggests that an increased curvature
- i. might exist.
l 3-8 Revised: 12-10-76
.mn_ -m.n.,a- m ,~ ,n n ,,n , , , .,,mm,, -, -- -
L ~
Subsequent to the removal of the delaninated cap (see Section 5.2.4) and prior to placing a new cap (see Section 5.2.8), a more extensive survey was conducted. Elevations were obtained at 2.5' horizontal intervals along the exposed outer conduit. Figure 3-27A is a plot of the-elevations for the West portion of tendon D-118. The elevations obtained and plotted are for the outer surface of the conduit. The prestress wires, of course, bear.against the structure and induce the prestressing force on the inside bottom of the conduit. Assuming the profile of the conduit inner surface is the same as the profile of the top outer. surface, tendon D-118 would tend to bear on the conduit at the high points and would tend to be straight between them. In the case of tendon D-118, the concentrated load at one of the bearing points based on an initial magnitude of the prestress force (1650 kips) would be approximately 180 kips. Although this value would be distributed along a finite length, it would result in a change in local load due to prestress for tendon D-118 from 41 psi to approximately 100 psi.
The survey data was also evaluated by fitting a circular curve to each data point using the adjacent points on either side of the point being considered. Average radii of all the outer conduit was computed to be 111.4' compared to the specified theoretical average radius of 110'.
However, the individual variations from the theoretical radii were significant. Figures 3-27B through 3-27E present the calculated radii
, for the outer tendon group.
i' The significance of the increase in radial tendon pressure due to either concentrated loads or decrease in tendon radius was analyzed using two approaches. The first studied the effect on radial tension of a ring load 4' wide, 40 feet from the apex and the second the effect of a 4' diameter uniform load at the apex. This increase of load in local areas resulted in the following: four (4) psi increase in load 1' results in one (1) psi increase in radial tension stress at th( level of the outer conduit. For the survey example discussed previously the 100 psi local pressure results in a nominal radial tension stress of 56 psi.
The survey and analysis results indicate that conditions that would have increased loads in local areas existed in the CR3 dome. This would have resulted in increased radial tension stresses in local regions of the dome.
8 J
. s 3-8a Revised: 12-10-76 mg.wwremm~ ww- "~w~ n w 'w ~:~~ ~~ " ~ ~
I 3.3.6 Heavy Construction L3 ads The construction of the dome did not require heavy construction loads to be placed on the dome. A small mobile crane was located on the ring girder for use in tendon installation and stressing, but no heavy equipment was located on the dome surface.
3.3.7 Coastal Location It was postulated that the coastal environment might have had a detrimental effect on the material in the structure. Subsequent investigation did not reveal detrimental effects due to coastal location.
3.3.8 Location Adjacent to Fossil Units There was a period of time when tendon conduits were exposed to the atmosphere prior to concreting. Although a potential for damage due to the presence of sodium, calcium, and magnesium sulfates from the stacks of the adjacent fossil units exists,(2) there is no evidence that reactions occurred.
3.3.9 Construction Methods Concrete construction joints were located as shown in Figure 2-3 and (V
the concrete placed in the sequence and on the dates ' indicated in Table 3-5. The field coring investigation program did not indicate unsatisfactory conditions such as honeycombing, voids or cold joints.
During construction a deficiency in concrete cylinder compressive ~
strength in pour "L" was observed. However, the concrete attained approximately 5800 psi at 90 days which was satisfactory (see Appendix B).
The specified prestressing sequence is shown in Figure 3-28 through 3-32 and given in Table 3-6. Three tendons, one from each layer or group, were to be stressed at a time (a sequence). Each tendon was to be stressed from both ends. The concept of the sequences was to put balanced loads into the dome.
G 3-9 Revised: 12-10-76
_ , , ;ge - . , -- -
-n -=-"= ~ =~ = r~:- - - -
l The stressing sequence log developed from constructor's records is given in Table 3-7. ' Table 3-8 shows the daily cumulative totals of tendons stressed. The stages of the stressing are shown in Figures 3-33.through 3-43. ~ Figures 3-44 and 3-45 show a relative vertical force imbalance on the done.~ The Tables and Figures indicate that at certain times the sequence'of stressing gave unbalanced forces as high as 16.9%. Immediately prior to the boom or loud noise on the morning of December 4, 1974,' 72% of the done tendons had been stressed. It is extremely-difficult to accurately assess the consequences of any given
, sequence but, the stressing sequence used did create unbalanced forces. ,
A field investigation was conducted by coring through the delsminated cap and using. survey methods to establish thicknesses of the dome and
, the delaninated cap and the location of the top of tendon conduit. The results are shown in Figures 3-24 through 3-27, and reflect concrete thicknees variations between 33-1/4 in, and 59 in, and top of conduit
, location from exterior concrete surface which varies between 6-1/2 in.
. and 13-1/2 in.
The effect of these variations is not considered significant. However, small force imbalances might have significance if coupled with tendon misalignment as discussed in Section 3.3.5.
3.3.10 Impact' Loads
, The boom heard by the workers at the site at 7:20 a.m. on December 4,1974 was discussed in Section 3.1.
, This boom might have been caused by the breakage of tendon wires or by a kink in a tendon i adjusting itself to the proper position. In either case, this boom 4
could have. represented an impact load applied to the dome causing initial cracks.
End caps were removed from each and of the tendons which were i .- detensioned or used for life off tests (see Section 3.1.3). Inspection of the anchorage assemblies did not revgal any conditions which might be associated with sudden breakage.
1
'3.3.11 Shrinkane Effects '
Because.of the-presence of the steel liner on the inner surface of the dome, the exterior surface concrete will shrink more rapidly than the
. balance of the concrete. The in-plane dimensional changes due to
, shrinkage produce meridional and hoop tension in top surface of the dome and meridional and hoop compression in the bottom surface. This effect should not produce radial tension near the top surface.
C f
3-10 Revised: 12-10-76 nmswwwymw ,-,-O mr-r- S~ ~ w - - mm m~-- n m~ w r ,e---'
i' During the three-month construction delay between April 4,1974 and July 8,1974 at construction joint L-M, shrinkage could have also affected the concrete. Since the concrete on the top surface and at the construction joint shrinks and the conduits protruding from the joint- do not shrink, micro-cracks could be produced in the concrete surrounding the conduits. This is similar to the case of micro-cracks in the mortar surrounding coarse aggregates due to shrinkage (l).
Figure 9 of Reference 1 indicates that when the distance between coarse aggregates is greater than 0.45 times the radius of the aggregates, shrinkage will cause radial compression, hence, hoop tension surrounding the coarse aggregates.
I f'
's f
C 3-10a Revised: 12-10-76 v enyyyem w-~rnv-~m- -~~~n~~- -- --
-m.~------
i i
In order to evaluate the amount of radial tension due to shrinkage for the three month construction delay, the amount of shrinkage has to be estimated first. The average final shrinkage of concrete (i.e. at 40 years with volume / surface ratio of 24 in. has been estimated to be 1
- x 10- in/in. However, based on the ASTM C 157 test of the design mix by the Pittsburgh Testing Laboratory, the shrinkage strain at 12 weeks is 4.5 x 10-4 in/in. Taking into account the difference between testing specimen and in-place concrete, it is reasonable to assume that
- shrinkage strain at the construction joint surface at the end of 12 weeks is 2.25 x 10-4 in/in Assuming thernal expansion coefficient of concrete to be 5.5 x 10-6 in/in/0F, the shrinkage strain' is equivalent to 410F temperature drop.
A 20" x 20" plane stress concrete analytical model with a ring sliced from 5 in, diameter schedule 40 pipe at the center as shown in Figure 3-22 was used tc evaluate the shrinkage stress. With a 41 F temperature drop applied at the surface of the concrete, the hoop tensile stress surrounding the conduit in the concrete is approximately 360 psi. The stress profile surrounding the conduit at the construction joint surface is shown in Figure 3-46. These stresses are based upon conservative shrinkage strains, but suggest sufficient magnitude to initiate cracking. Assuming a shrinkage strain of 2.25 x 10-5 in/in for the overall dome, the tensile stress surrounding the conduit is 36 psi.
3.4 CONCLUSION
S
(
It appears that a compression-tension interaction failure occurred.
Effects which could have generated radial tension forces have 'seen defined and discussed. Several of the effects such as radial tension
, due to pre'tressing, s thermal effects, tendon alignment, stress concentrations and shrinkage in combination would have been sufficient when combined with biaxial compressive stresses and lower than normal dir'ect tensile strength of the concrete to result in the delaminations.
The complete fracture of the coarse aggregate on that surface and the variations in tensile strength values obtained from the direct tensile tests indicate that the fragility of the coarse aggregate permitted local cracking to propagate.
(.
3-11 Revised: 12-10-76 A '"
WQQY: e lW ?#*} Yf *' # ~'?"# ? '"" '
?' : r e-
/ .
TABLE 3-1 INITIAL EXPLORATORY CORE HOLES Location Delamination Core No. Dias. (in) Az Pour Thickness (in) Gap (in) 1 1-3/4 -90* L 11-1/2 1-1/2 2 4 270 0 L 13-3/8 1-5/8 3 4 90' Q 15-1/4 1-3/4 4 4 270* J 4-1/2 1/8 5 4 150 L 14 2 6 4 30 L 10-1/2 1-1/2 7 4 335 ~ L 13 1-3/4 t
g l
12 -l
=rgmus~ a ~. -
TABLE 3-2 CORE STRENGTH TEST RESULTS Compression Compression Test Moving Average Split Pour Core Strength Average of Three Tension Remarks G(W) 49 7000 - - -
G(W) 50 6820 6910 - -
G(W) 51 - - -
G(E) 55 5690 - - -
G(E) 56 6330 6010 - -
G(E) 57 - - - -
, G(N) 58 7080 - - -
G(N) 59 6160 6620 6513 -
._ G(N) 60 - - -
705 G(S) 52 6680 - - -
G(S) -5. 5970 6320 6316 -
G(S) 54 - - -
625 H(SE) 45 6580 - - -
. H(SE) 13 5330 5950 6296 -
H(SE) 46 - - -
710 j H(NW) 40 6860 - - -
H(NW) 41 6180 6520 6263 -
i i
H(NW) 42 - - - -
H(SW) 43 6620 - - -
H(SW) 12 6010 6310 6260 -
, H(SW) 44 - - -
730 3-13 q egy#~ 3. -~-.& .-
TABLE 3-2 (Cont'd)
Compression Compression Test Moving Average Split Pour Core Strength Average of Three Tension Remarks H(NE) 47 5440 - - -
3(NE) 48 7090 6260 6363 -
H(NE) 11 - - - -
J(5) 34 6990 - - -
J(5) 35 6620 6800 6456 -
J(5) 36 - - - -
J(6) 37 5220- - - -
J(6) 38 6440 5830 6297 .
1 J(6) 39 - - -
720 K(7) 31 6920 - - -
K(7) 32 6480 6700 6443 -
K(7) 33 - - -
780 K(7) 10 6510 - - -
Specimen test dry - j not included.
K(8) 28 5940 - - -
K(8) 29 6090 6010 6180 -
K(8). 30 - - - -
L(9) 9 6250 - - -
Specimen tested dry -
not included.
L(9) 25 4040 - -
Flaw in specimen -
not included.
L(9) 26 6560 - - -
L(9) 61 5900 6230 6313 -
L(9) 27 - - -
625 3-14
. p, ,_g., ,.c 7- - ---
b TABLE 3-2 (. Cont'd)
Cnoression Compression Test Moving Average Split Pour Core Strength Average of Three Tension Remarks M(10) 22 5890 - - -
M(10) 23 6610 6250 6163 -
M(10) 24 - - - -
N(11) 11 - - - -
N(11) 19 5370 - - -
N(11) 20 6650 6010 6163 -
N(11) 21 - - -
675 P(12) 16 6360 - - -
P(12) 17 6090 6220 6160 -
P(12) 18 - - - -
Q(13) 8 6500 - - -
Q(13) 14 5760 6130 6120 -
Q(13) 15 - - -
800 Notes
- 1. Number of compression test specimens = 34
- 2. Lowest average of three compression tests = 6130 psi
- 3. Lowest set of compression specimens + 500 psi = 5830 + 500 = 6330
- 4. Average compression stress = 6301 psi S. Number'of tension test specimens = 9
- 6. Average strength of tension test specimens = 708 psi 3-15
- tw, N, . Fu ;
L f
- )
N 3
TABLE 3-3
- , CORE IDCATIONS AND MAIN DEIAMINATION DETAILS t
r -
Cap Cap Core i Azimuth Radius Dimension '
Thickness Remarks 1 92 16'30" 31'-10-1/4" 1-1/2" 11-1/2" 1-3/4" core 2 264022'00" 32'-2-3/4" 1-5/8"- 12-7/8" 3 97 37'00" 3'-10-1/2" 1-3/4" 15'-4" 4B 2700 12'15" 50'-0" 1/8" 4-1/8" 5 151 36'00" 30'-10" 2" 13-1/2" 6 2601 6'00" 32'-2' 1-1/2" 10-1/8" 7 333 11'30" 31'-10-3/8" 1-3/4" 13" 8A 6 06'00" 3'-9-3/4"- 1-7/8" 14-7/8" 0B 3580 20'00" 3'-7-3/4" 2" 14-1/2" 9 3 58'45" 31'-3" 1-5/8" 14-1/8" 10 3 32'00" 36'-5-3/4" 1-1/2" 12-7/8" 11 359 54'45" 50'-4-1/2" 1/8" 5-3/8" 128 1190 38'30" 48'-11" 1/8" 4-5/8" Y 13 238026'15" 48'-6-1/4" 3/8" '
4-1/8" 5 14C 226 18'45" 3'-11-3/4" 1-5/8" 15" 15 282048'00" 8'-5-1/2" 2" 14-1/8" i
16B 42 56'15" 15'-2-1/2" 1-7/8" 13-1/4" 17 '186 08'00" 14'-3" 2-1/8" 13" 18 323026'45" 14'-4-3/4" 2" 14-3/4" 19 255" n'00" 19'-3-1/2" 2" 13-5/8" 20 115% 7'30" 19'-4-7/8" 1-5/8" 13-3/4" 21 246 22'45" 19'-3" 2" 12-1/8" 22 42054'0's" 25'-5-1/2" 1-3/4" 13" 23 180 42'00" 26'-10-1/2" 1-3/4" 14-1/8" 24 326 56'15" 26'-0" 2-1/8" 12-3/4" 25 1160 21'45" 30'-9-3/8" 2-1,' 9" 9-1/8" 26 230 15'45" 32'-1-1/2" 1-3/4" 11-3/8" 27 3290 21'00" 32'-9-1/2" 1-7/8" 13-3/8" 28 349 02'45" 36'-6-3/8" 1-3/8" 13-3/4" 29 37 58'15" 39'-4-3/4" 1-3/4" 10-3/8" 30 102 29'00" 38'-7" 1-3/8" 8-3/4"
3 y
e Y
TABLE 3-3 (Cont'd)
[tv Cap Cap Core # Azimuth Radius Dimension Thickness Remarks
'31 157 20'00" 39'-6" 1" 13-5/8" 32A 223*02'75" 38'-8-3/4" 1" 13-3/4" 33 3060 19'45" 36'-6-7/8" 1" 13-1/4" 34D 34 52'30" 43'-10" - -
ND 35 1310 04'45" 43'-4" 1" 10-7/8" 36B 210 57'00" 45'-1-1/2" 1/4" 7-3/4" 37 249 24'30" 44'-11-1/2" 1/2" 6-1/8" 38 318*04'00" 43'-3-1/2" 5/8" 12-1/2" 39 7 46'00" 43'-9-1/4" 3/8" 12-5/8" 0 0 408 71845" 49'-7-1/4" - -
ND 41B 23 04'30" 49'-9-1/2" - -
ND 42 78 55'30" 49'-5" 1/2" 2-1/2" 42A 76 04'45" 49'-5-1/4" 1/2" 4-1/8" l-3/4" dia. core 42B 78*56'00" 46'-1-1/2" 1/8" 6-1/8" 1-3/4" dia. core w 42C 79 00'00" 50'-3-1/2" 1/4" 2-1/8" 1-3/4" dia, core
,L 42D 810 55'30" 49'-2-3/4" 1/2" 3-7/8" 1-3/4" dia, core
" 43 97 28'15" 49'-5" 1/8" 1-3/4" 44 170 53'15" 48'-9-1/4" 1/8" 4-1/2" 45 191 10'15" 49'-8-1/2" - -
ND 46 259 54'00" 47'-C-1/4" - - ND 47B 280 24'30" 48'-7-1/2" - - ND 48C 311 54'00" 51'-9-1/8" - -
ND 49 37 10'00" 55'-8-1/2" - -
ND 50 90 58'00" 55'-10" - -
ND 51 122 46'30" 53'-1-3/4" - -
ND 52A '146 10'00" 52'-11" - -
ND 53 1830 15'00" 54'-11-1/2" - - ND 54D 218 56'45" 55'-4-1/4" - -
ND 55 237 23'45" 52'-10-1/2" - -
ND 56 268 29'15" 55'-8-7/8" - - ND 307 08'00" 55'-8" 57 - -
ND 58 335 31'00" 56'-0-1/2" - -
ND 59D 351 06'45" 55'-1-1/2" - - . ND 60 24 51'45" 57'-1-1/4" - -
ND 61(VII) 83 49'00" 32'-0-1/8" 1-1/2" 10-7/8"
4 ; I t
..s 4
- 3
' @ TABLE 3-3.(Cont'd) 6 Cap Cap Core i Azimutil Radius Dimension ' Thickness Remarks 9 I 73 49'45" 12'-3" 2" 13" Construction joint examination
~
IIIA 226 06'45" 24'-0" 1-1/2" ' 1/2" Construction joint examination 0
'IIIB 63g9'00" 23'-11-1/2" 1-1/2" 13" Construction joint examination IV 254 15'45" 30'-0" 2" 12-3/4" Construction joint examination VB 1601 8'30" 35'-11-3/8" 1-1/4" 13-1/2" Construction joint examination VI 4 18045" 41'-6-1/2" 1"- 12-1/2" Construction joint examination VIII 156 25'45" 13'-0-5/8" - -
ND - Tensile test core IX 336046'30" 15'-0-5/8" - -
ND - Tensile test core XII' 63 38'00" 19'-1-5/8" 2" 13-5/8" Tensile test core XV 160 06'00" 32'-11-1/2" 1-5/8" 10-3/4" Tensile test core XVI 10 23'00" 20'-11-1/8" 2" 12-7/8" Tensile test core XVIII 1130 21'15" 27'-8-3/4" 1-1/8" 13" Tensile test core 0 00'00" 20'-0" 2-1/8" 11-1/2" 1-3/4" tendon survey core 900 00'00" 20'-0" 1-3/8" 13-5/8" 1-3/4" tendon survey core y -
180 00'00" 20'-0" 2-3/8" 11-3/4" 1-3/4" tendon survey core g -
270 00'00" 20'-0" 2" 11-1/8" 1-3/4" tendon survey core 0000'00" 40'-0" 1" 10-1/4" 1-3/4" tendon survey core 900 00'00" 40'-0" 2" 8-3/8" 1-3/4" t'endon survey core 180 00'00" 40'-0" 1-1/2" 11-1/4" 1-3/4" tendon survey core
'270 00'00" 40'-0" 1-1/4" 10-1/8" 1-3/4" tendon survey c.te 900 00'00" 50'-0" 5/8" 2-7/8" 1-3/4" tendon survey core 180 00'00" 50'-0" 1/16" 4-7/8" . 1-3/4" tendon survey core Notes: a. Cap dimensions and cap thickness are radial measurements.
- b. ND denotes no delanination.
TABLE 3-3A REACTOR BUILDING DOME SECONDARY CRACK LOCATIONS
- D1 33 3/4" deep, 19 1/2" x hairline, 13" x 1" D2 32 1/2" deep, 21 3/8" x hairline, 14 1/2" x 1 1/2" D3 32 1/4" deep, 26 3/8" x 1/16", 23" x hairline, 19 1/2" x 1/4" **,
12 3/4" x 2 1/8" D4 31" deep, 24 1/4" x hairline, 22 1/16" x hairline, -12" x hairline, 9 3/16" x 3/8", 6" x 11/16" D5 -19 5/8" deep, hole not deep enough D6 32 1/8" deep, 25 3/4" x 1/32", 17 15/16" x 1/32", 17 13/16" x 1/32",
13 5/8" x 5/16", 10" x 1 1/8", 6 7/16" x 3/16"
.D7 32 1/2" deep, 19 1/2" haitc.ne on North half, 13 5/8" x hairline, ;
, 12 1/4" x 5/8" !
- [' D8 32 1/4" deep, 27 1/4" x 1/32", 15 13/16" x 5/16" **, 12 7/16" x 1 11/16",
10 3/4" x 1/8", 9 1/4" x 3/8" **, 8 1/2" x 1/2" **, 8 1/16" x 3/16" **
D9 32 1/4" deep, 28 3/4" x 1/32", 19 1/2" x 1/8", 16 5/8" x 3/8",
13" x 2" with suspended slice of concrete, 9 5/8" x 3/16" **
D 10 33 3/8" deep, 25" x hairline, 22 1/ 2" x 1/32",14 3/4" x 2 1/4" D 11 31 1/2" deep, 25 1/2" hairline, 21 5/8" x 1/16", 12" x 2" D 12 34 3/4" deep, 25 7/8" x hairline, 22 3/4" x 1/16", 13 5/8" x 2 1/8" D 13 37 1/2" deep, 19" x 1/16", 17 1/2" x 1/16", 14" x 2" D 14 34 1/2" deep, 23 5/16" x hairline, 13 7/16" n 1 7/8" D 15 33" deep, 25" x 1/16", 20" x hairline, 17 1/16" x 7/16", 13 3/4" x 2 D 16 32 3/4" deep, 26 1/4" x hairline, 16" x ** in 1/8" crack, 9 11/16" x 13/16" D 17 38 1/2" deep, 33 3/8" x hairline, 28 1/2" x hairline, 19 1/4" x 1/16" **,
-19 3/8" x 1/16" **, 15" x 2" l
- For plan location see Figure 3-10D.
- Multiple delamination indicated.
3-18a Revised: 8-10-76 m ww.m wc' s
~ -
r- ~ ~ -- --~- = - - -
t.
- TABLE 3-3A (CONT'D)
D 18 34" deep, 27-3/8" x hairline, 18 1/4" x 1/8" **, 13 5/8" x 2" D 19 34 1/4" deep, 29'1/8" x' hairline, 16" x hairline, 13 1/4" x 2" D 20 34 1/4" deep, 23 1/4" x 1/32", 17 5/8" x 1/16", 11 11/16" x 1 3/4" D 21 34" deep, 26 7/16" x hairline, 23 3/8" x 1/32", 20 3/8" x 1/8",
14 1/8" x 2 D 22 35 3/8" deep, 25 9/16" x hairline, 19 1/4" x ha'irline, 12 3/16" x 1 9/16" D 23 32 1/4" deep, 16 11/16" x 3/32", 11 1/8" x 1", 10" x 3/8"
. D 24 34 5/8" deep, 9 3/4" x 3/8", 9" x 1/32"
- Multiple delamination indicated.
3-18b Revised: 8-10-76 rwar:wse;w.m - h: , - :-:-'~'m- ~ ' - * ~ -
-'--~ ~ - ~ ~ ~ - - ~ ' - ~
TABLE 3-3B i AVERAGE LIFT-OFF DATA Tendon Date Ave.
No. Tensioned Liftoff Remarks D108 12/9/74 1525 ^
Dil4 12/11/74 1450 D120 12/5/74 1448 D121 11/19/74 1342 D122 12/2/74 1498 D127 12/5/74 1465 D128 11/5/74 1355 D129 11/19/74 1465 D134 12/3/74 1515 D137 12/10/74 1418 D201* 11/14/74 (12/9/74) 1535 l D208 12/4/74 1500 D214 12/6/74 1435 Tendon was not detensioned D221 12/12/74 1490 D228 12/2/74 1368 D234 12/5/74 1535 D308 12/6/74 1475 D314 11/27/74 1453 D321 11/20/74 1545 D318 2/5/74 1523 D334 12/4/74 1333 y
~ ^
D101 11/26/74 1553 D109 11/21/74 1485 D117 11/21/74 1510 D125 11/19/74 1373 D133 11/18/74 1553 D141 11/15/74 1577 D201 11/14/74 (12/9/74) 1485 l D209 12/5/74 1523 D217 11/19/74 1520 Tendon detensioned D225 11/20/74 1420
-D233 11/22/74 1618 D241 11/25/74 (4/2/75) 1488 l D301 11/25/74 1583 D309 11/22/74 1577 D21s 11/21/74 1546 D325 11/19/74 1593 D333 11/18/74 1600 D341 11/15/74 1474 y Average 1491 Predicted Design Value = 141K/IN x 9.72.IN / TENDON = 1370K *
- Denotes tendon subsequently detensioned .
i Dates in parentheses are for final retensioning, see Table 3-7 3-18e Revised: 12-10-76
) F [4 ' uO j
g y. .!' \
f'Y
- TABLE 3-4 zi- DOME EP0XY AND CONCRETE STRENGTH TESTS N
N ~
Pour Epoxy 28 Day Concrete 90 Day Concrete Moving ' Moving y Avg. to Avg / Average Avg /- ' Average Date No. PSI Date PSI Fet of Thre.4 PSI Set of Threc
~!
6560 7060 I 2/18/74 976-RB 6190 -6580 (G-la) 6340 7831 6470 6312 6830 .6622 6030 5820 2/20/74 978-RB 5240 7776 6330 6510 6370 6695
_(G-lb) 6690 7029 2/25/74 780-RB 6320 7746 4600 5255 6026 ' 7420 7175 6830 (G-2a) -5910 6930
,L 2/25/74 981-RB D.O. D.O. 5560 5785 5850 6720 5790 6553
- 6010 (G-2b) 4860 5280 5980 3/1/74 783-RB 5680 6380 (H-3a) 8590 7763 4820 5102 5381 5860 5762 6242 4630 5380 3/4/74 985-RB 9190 7791 5540 5725 5537 6770 6780 6110 (ll-3b) 5910 5790 3/6/74 987-RB 8680 7809 6050 5450 5426 6280 6220- 6254 (ll-4a) 4850 6160 3/7/74 988-RB 8425 7821 ' 5.800 5560 5578 6070 6035 6345 (11-4b) 5320 6000 6540 6380 3/12/74 992-RB 8000 7824 5870 6078 5696 6580 6580 6278 (J-5) 5430 6790 6470 6580
y ra f %,
E m
-14 h
, TABLE 3-4 (Cont'd)-
y
-N* Pour Epoxy 28 Day Concrete '90 Day Concrete fj Moving Moving-p Avg. to Avg / Average Avg / Average j Date No. PSI Date PSI Set of Three PSI Set of Three 3/13/74 994-RB* 5310 6000 (H-2a) (10,175) -
4790 4940 5526 5610 6050 6221 4880 6190 4780 6400
, 3/14/74 995-RB 5570 7590 (J-6) 9840 7862 5840. 5868 5629 6190 6835 6488 3 6470 6490 5590 7070 3/18/74 996-RB 4980 5960 6 (R-2b) (9170) -
5040 5028 5279 6150 5928 6271' E 5180 5360 4970 6240 3/20/74 997-RB .4420 6420 (K-7) 8530 -7874 5110 4918 5271 6930 6430 6397 5360 6240 4780 6130 3/22/74 998-RB 5090 6700 (R-la) (5220) -
4560 4978 4984 6460 6322 6226 6360 6240 4900 5890 3/22/74 999-RB 4950 5390 (R-lb) 0.0. 4810 4880 4925 5870 5630 6127
- Pour location can not be identified.
1 b
)
. .I-TABLE 3-4 (Cont'd) y Pour Epoxy 28 Day Concrete 90 Day Concrete
' .. Moving Moving
'$'J Date Avg. to Avg / Average Avg / Average No. PSI Date PSI Set of Three PSI Set of Three 3/26/74 1000-RB 5410 6440 (K-8) 7370 7865 4690 5305- 5054 6370 6338 6096 2 5620 5840 5500 6700 4'860 4920 4920 6620 q 4/4/74 1004-RB 4580 5960
! (L-9) 5560 7825 4560 4571 4919 -
5820 5929 4320 5780 4240 -
Y 7040 6720 0 6980 7420 7/8/74 1009-RB 6400 7360 (M-10) 7200 7814 6050 6583 5486 7420 7185 6447 6600 6920 6420 7070 6560 .
7070 7/12/74 1010-RB 6760 7410 (N-ll) 5530 7775 6540 6618 5924 7270 7175 6726 6610 6950 6560 6930 7/17/74 1012-RB 6670 7040 (P-12) 7170 7765 6230 6342 6514 6690 6770 7043 5910 6420 6560 7/22/74 1014-RB 6440 (Q-13) 7475 7760 . 6790 6670 6543 6970 i
TABLE 3-5 DOME POUR LOG Pour No. Date Location 1
-G1 2-18-74 45 - 135 G1 2-20-74 225 - 315 G2 2-25-74 135 - 225 G2 2-25-74 315 - 45 H3 3-1-74 180 - 270 H3 3-4-74 0 - 90 H4 3-6-74 90 - 180 H4 3-7-74 270 - 360 JS 3-12-74 45 - 225
. J6 3-14-74 225 - 45 K7 3-20-74 135 - 315 K8 3-26-74 315 - 135 L9 4-4-74 0 - 360 M10 7-8-74 0 - 360
. N11 7-12-74 0* - 360 P12 7-17-74 0 - 360 4 Q13. 7-22-74 0 - 360
,e 3-22
' se+wm, pre:: -- - g-- ,
-m- --- -
TABLE 3-6 PRESTRESSING SEQUENCES Sequence Sequence No. Tendons No. Tendons
- 1. D131 D211 D331 22. D102 D240 D302
- 2. D127 D215 D327 23. D106 D236 D306
- 3. D123 D219 D323 24. D110 D232 D310
- 4. D119 D223 D319 25. D114 D228 D314
- 5. -D115 D227 D315 26. D118 D224 D318
- 6. D111 D231 D311 27. D122 D220 D322
- 7. D107 D235 D307 28. D126 D216 D326
- 8. D103 D239 D303 29. D130 D212 D330
- 9. D135 D207 D335 30. D134 D208 D334
. 10. D139 D203 D339 31. D138 D204 D338
- 11. D141 D201 D341 32. D140 D202 D340
- 12. -D137 D205 D337 33. D136 D206 D336
- 13. D133 D209 D333 34. D132 D210 D332
- 14. D129 D213 D329 35. D128 D214 D328
, 15. D125 D217 D325 36. D124 D218 D324
- 16. D121 D221 D321 37. D120 D222 D320
- 17. D11' D225 D317 38. D116 D226 D316
- 18. D1; I D229 D313 39. D112 D230 D312
- 19. D109 D233 D309 40. D108 D234 D308
- 20. D105 D237 D305 41. D104 D238 D304
- 21. D101 D241 D301
.A 3-23 Revised: 12-10-76
~ svrm w . emw-~ e- -
= , - - - ---m-~-- - -- - - - - - -
Ec IABLE 3-6 PRESTRESSING SEQUENCES Sequence Sequence No. Tendons No. Tendons 1.- D131 D411 D331 22. D102 D240 D302
.2.
D127 D215 D327 23. D106 D236 D306 3, D123 D219 D323 24. D110 D232 D310
'4. D119 D223 D319 25. D114 D223 D314
- 5. D115- D227 D315 26. D118 D224 D318
- 6. Dill D231 D311 27. D122 D220 D322
- 7. D107 D235 D307 28. D126 D216 D326
- 8. D103 D239 D303 29. D130 D212 D330
~
- 9. D135 D207 D335 30. D134 D208 D334
- 10. D139 D203 D339 31. D133 D204 D338
- 11. D141 D201 D341 32. D140 D202 D340
-12. - D137 D205 D337 33. D136 D206 D336
- 13. D133 D209 D333 34. D132 D210 D332
- 14. D129 D213 D329 35. D128 D214 D328
- 15. D125 D217 D325 36. D124 D218 D324
- 16. D121 D221 D321 37. D120 D222 D320
- 17. D117 D225 D317 38. D116 D226 D316
- 18. D113 D229 D313 39. D112 D230 D312
- 19. D109 D233 D309 40. D108 D234 D308
- 20. n105 D237 D305 41. . D104 D238 D304
- 21. D101 D241 D301 a..
3-23 7pr.ww? < c.>1 mn ::,;u , ,
1
TABLE 3-7 ACTUAL TENDON STRESSING LOG C
Date Tendons 10-25-74 D331 10-30-74 D211 10-31-74 D130 11-4-74 D215 11-5-74 D128 11-7-74 D327, D219 11-8-74 D323, D123, D319 11-11-74 D223, D115, D227, D315, D111, D311, D119 11-12-74 D231, D107, D235, D307, D239*
11-13-74 D103*, D102*
11-14-74 D207, D303, D135, D201*, D335 11-15-74 D139, D339*, D341, D203, D141 11-18-74 D204, D338, D137*, D133, D210 D333, D213 11-19-74 D329, D129, D125, D325, D217, D121 11-20-74 D221*, D225. D321
, 11-21-74 D117, D317, D113, D313, D229, D109
(_ 11-22-74 D233, D309, D105
~
11-25-74 D238, D241*, D305, D301 l 11-26-74 D302, D104, D101, D240, D236, D232 11-27-74 D306, D310, D314, D106, D110, D114* l 12-2-74 D118, D122, D318, D228, D322, D326 12-3-74 .D224, D330, D216, D220, D132, D212, D126, D134 12-4-74 D138, D208, D334, D336, D337, D340, D140, D205, D222, D226 12-5-74 D230, D332, D234, D328, D237, D136, D131, D127, D209, D124, D120, D116 12-6-74 D206, D320, D324, D316, D312, D308, D304, D202, D214, D218 12-9-74 D112, D108, D201, D103 12-10-74 D137 12-11-74 D114, D102 12-12-74 D221 3-31-75 D339 4-2-75 D239, D241
- Denotes tendons which were detensioned and retensioned at a later date to satisfy Nonconformance Reports.
~
3-24 Revised: 12-10-76
-ymum ~~;. . v -
~e w - - - - -- --
n
-TABLE 3-8 TENDON STRESSING / CUMULATIVE DAILY TOTALS Layer Layer Layer Date 1 2 3 10-28-74 0/1 0/0 1/1 10-30-74 0/0 1/1 0/1 10-31-74 1/1 0/1 0/1 11-1-74 1/2 . 0/1 0/1 l'l-4-74 0/2 1/2 0/1 11-7-74 0/2 1/3 1/2 11-8-74 2/4 0/3 2/4 11-11-74 2/6 2/5 2/6 11-12-74 1/7 3/8 1/7 11-13-74 2/9 0/8 0/7
, _ . 11-14-74 1/10 2/10 2/9 11-15-74 2/12 1/11- 2/11 18-74 2/14 3/14 2/13 11-19-74 3/17 1/15 2/15 11-20-74 0/17 2/17 1/16 11-21-74 '3/20 1/18 2/18 11-?2-74 1/21 1/19 1/19 11-25-74 0/21 2/21 2/21 11-26-74 2/23 3/24 1/22
.11-27-74 3/26 0/24 3/25 12-2-74 2/28 1/25 3/28
~
3-74 3/31 4/29 1/29 12-4-74 2/33 4/33 4/33
,_ 12-5-74 6/39 4/37 2/35 12-6-74 0/39 4/41 6/41 12-9-74 2/41 0/41 0/41 3-25 ggMwymemnom ~ en - "" ~
":' I ,_. - .-
f'f' a l(m 4.0 DELAMINATED STRUCTURE
4.1 INTRODUCTION
-The delsminated dome condition was evaluated for the various critical l design bases assuming an effective 24" thickness except in the proximity
-of the ring girder where field investigations revealed an approximate l thickness of 36". The resulting analytical model consisted of a shell of symmetric configuration, theoretical radii and the aforementioned effective concrete thicknesses -(i.e. deep cracks were not represented).
-The dome.model's geometry considered in the analyses is shown in Figure 4-1.
4.2 APPLICABLE CODES AND STANDARDS For the evaluation of the delaminated dome the following documents were considered:
- 1. Crystal River Unit 3 FSAR Docket No. 50-302.
- 2. ASME Boiler and Pressure Vessel Code,Section III, Division 2, 1975. (liner strains only) 4.3 CRITERIA
- I
, The applicable criteria are given in Table 2-2.
4.4 EVALUATION .
The objet.ive of the analyses of the delaminated structure was twofold:
- 1. to confirm hand calculations related to the safety of the structure,
- 2. -to investigate the serviceability of the structure without major modification.
The load combinations investigated with the exception of the small break accident are given in Table 2-3. The evaluation which follows was based on an in-place strength of concrete of 6000 psi (refer to Section 3.3.1).
The results are presented for the delaminated structure using FSAR and
- i. ASME Code Section III Division 2 (liner only)' acceptance criteria.
4 These criteria as they apply to-extreme fiber compression stresses, membrane stresses and liner strains are given for each load combination.
4 The acceptability of the dose for shear was evaluated by the following l procedures: .
The minimum capacity'of the concrete, $ Vc, was compared with the l ultimate shear, Vu, that exists on the section. If a case results
, where this minimum concrete capacity was less than the ultimate shear, then the~ actual section capacity was computed using the
_ FSAR criteria.
. 4-1 Revised: 12-10-76 semkswrpm& : n:~n- m- v -- . - - - --- -
l The shear capacities described above represent ultimate capacities.
For. assessing the section adequacy in shear, the ultimate shear Vu, was l calculated by applying a load factor of 1.5 to the net shears resulting from load combination a, b, and e in Table 2-3. A load factor of 1.0 was applied to the net shears for load combination d. l The shear provisions of section 2.3.2 were applied in the delaminated dome shear investigations.
The location of the stations in degrees noted en the shear figures is shown in Figure.4-1.
The consideration of radial tension in combination with radial shear is discussed in Appendix I.
The structural analysis of the containment was performed using KALNIN'S Static Computer Program described in Appendix D. The individual loads which comprise the load combinations were input separately, and their results were combined internally in the program where possible. This was not possible for the Structural Integrity Test and Accident Condition load combinations due to the different Young's Modulus (E) values for the concrete under the sustained loads (D, F, and T )o and the rapidly applied loads (Pa and T,). Ir,these cases, stresses for each of the two types of loads were combined externally to the computer. The effects of shrinkage and creep were considered as discussed below. .
- a. Shrinkage -
The effect of concrete shrinkage on the overall structural response (stress resultants) was insignificant due to the large l volume to surface ratios of the cylindrical wall, ring girder, and dome.
, In the prestress loss calculations a conservative value for long term shrinkage strain of 100 micro in/in was used to be l consistent with the original design. This is the value recommended in Reference (12) for calculating prestress losses in f' = 5000 psi concrete. Actually, use of the shrinkage equation appearing in this reference for time = 40 yr. and volume to surface ratio = 24" results in a shrinkage strain of 10 micro in/in at end of plant life.
- b. Creep The effect of concrete creep under the prestress loads was included in the prestress loss calculations and in the structural analysis.
The creep curves appearing in Reference (12) allow specific creep strains to be determined'considering both concrete age at loading and duration of load. Actual creep strains were calculated from ,
these specific creep strains for use in determining prestress !
losses. Also, the reduction in concrete stresses, which results I in an increase in liner stresses, caused by concrete creep unt r l i
l l
4-2 Revised: 12-10-76 l
m m nmmv m-= -- - -
sustained loads was taken into account in the structural analysis
' by using an effective Young's Modulus, E'. This modulus is expressed in terms of specific creep as c
E c " 1 + sc E '
where:
0 Ec = instantaneous concrete Youag's Modulus = 4 x 10 psi, sc = specific creep (micro in/in/ psi)
Analysis of the containment for load combinations a, b, and c (Table 2-2) was based on calculated prestress losses and a sustained l load (D F, To ) Ej = 2.7 x 10 6psi corresponding to the present time. In load combination c (SIT), the results for 1.15P were l a
based on E = 4.0 x 106 pai, For investintion of the containment under load combination d(LOCA),
40 year calculated values of prestress losses and Ej = 1.8 x 106 psi were seed. The 1.5P and T parts were based on E =4x l a a c 106 psi.
[ c. Prestress Losses -
The calculated prestress losses (ksi) and effective prestress (ksi) are given below:
Elastic Steel Total Effective Shortening Creep Relaxation Shrinkage Losses Prestress Present Vertical 3.6 3.9 2.2 2.9 12.6 155.4 Hoop 6.4 7.0 2.2 2.9 18.5 146.2 Dome 10.4 11.3 2.2 2.9 26.8 141.2 40 vr.
Vertical 3.6 9.1 3.4 2.9 19.0 149.0 Hoop 6.4 16.2 3.3 2.9 28.8 136.0
' Dome 10.4 26.2 3.4 2.9 42.9 125.1 l
I l
s 4-3 Revised: 12-10-76 spympwm. ~,., .- - 7v - - ~ ~ .
~---m--- - - - - -
(- The membrane and extreme fiber stress results presented in this section were those obtained directly from the KALNIN'S Static Computer Program l analyses (i.e. linear, elastic, uncracked). At locations in the dome where tensile stresses exceeded the allowable values given in Table 2-2, the concrete was assumed to be cracked. Cracked section investigations were performed to calculate concrete compressive and rebar tensile stresses. In the cracking investigation, the axial force (P) and moment ('f) stress resultants applied on the section were computed from the uncracked stresses (plotted). The only exception to this was for the Normal Winter Operating Condition load combination. In this case, the cracked section reduccd the effect of the through thickness gradient l (AT) part of the To term in the load combination. Therefore, the uncracked stresses due to AT were subtracted from the plotted stresses l prior to computing P and M. Then, the effect of AT applied to the section with P and M was considered in a manner similar to that described l in ACI 505-54. Cracked section stresses, calculated as described above, are shown at selected locations in some of the figures for this section.
4.4.1 Structure Prior to Operation For this load combination, the allowable extreme fiber stress according a to the FSAR is 0.6 f' compression and zero tension. .This results in an allowable stress of 3600 psi compression for 6000 psi concrete. For an uncracked section, the results in Figure 4-2 indicate that the
[, stresses near the ring girder are tensile. Thus, a cracked section investigation was required. This resulted in a peak compressive stress l of 3326 psi. Figure 4-3 shows a peak compressive hoop stress of 2413 psi.
No hoop tensile stresses exist.
The allowable membrane compression stress using tha.FSAR is 0.45 f'.
This results in an allowable compressive stress of 2700 psi for 6000 psi concrete. Results shown in Figure 4-4 indicate a maximum membrane compression stress of 2470 psi. No membrane tension exists.
The ASME Code limits for liner strain for this load combination are a compression strain of 0.002 and tension strain of 0.001. Figure 4-5 shows a maximum compression strain of 0.0011 in the meridional direction and a maximum tensile strain of 0.000002 in the hoop direction.
The shear stress limits noted in the FSAR were used in this evaluation. l Figure 4-6 shows that the available shear capacity is in excess of the required shear capacity. Thus the delaminated structure was considered j serviceable for this load combination.
4-4 P.evised: 12-10-76 w w n n w :~ r m ~ -- - - - -- --
'k Figure 4-12 shows that the available shear capacity is in excess of the actual shears using the FSAR criteria.
4.4.3- Structural Integrity Test The allowable. extreme fiber stresses according to the FSAR are 0.6 f' compression and zero in tension. This results in an allowable l stress of 3600 psi compression for 6000 psi concrete. The uncracked l results shown in Figures 4-14 and 4-15 show a peak compressive stress ;
of 1655 psi and a peak tensile stress of 186 psi. Based on a cracked section investigation, a concrete compressive stress of 1682 psi and compressive stress in the rebar exists at this location since the l cracking does not extend far enough into the section to reach the inside i face rebar. '
1 For the condition of membrane stress the FSAR limit is 0.45 f' compression I and zero tension. This results in an allowable compressive stress of 1 2700 psi for 6000 psi concrete. Figure 4-16 shows the peak membrane compressive stress to be 902 psi, which is below allowable values. No l ,
membrane tension existe.
Liner strain limits given in the ASME Code are a compressive strain of 0.002 and a tensile strain of 0.001. Figure 4-17 shows a peak liner compressive strain of 0.00057. For this condition the analysis indicates no liner tensile strain.
Using FSAR criteria, Figure 4-18 shows the available shear capacity exceeds the required capacity.
4.4.4 Accident Condition The delaminated structure was investigated for a late plant life (40 l years) loss of coolant accident condition. The 40 year prestress losses, accompanied by a sustained load reduced concrete modulus, Ee ,
were considered. The accident temperature effect on the liner up to 36 ksi yield produces a tensile force in the concrete and was included.
The results for an uncracked analysis are shown in Figures 4-20 and n*
4-21.
The membrane stresses illustrated in Figure 4-20 are tensile over most of the dome. In accordance with the Design Criteria, Table 2-2, the concrete is assumed to have zero tensile capacity. Since the delaminated I dome does not have sufficient reinforcement to resist these membrane tension forces, a cracked analysis of the dome was performed relying on the resistance-of the unbonded tendon system.
For the accident suudition, the tendon stress is limited by the ASME Code to 0.9 f py, whichEis equivalent to 0,.72 f pu. The allowable liner strains f rom the ASME Code ~ are 0.005 in compression and 0.003 in tension. This analysis did~not consider the 5 inch schedule 40 pipe as reinforcement.
x
}
4-5 Revised: 12-10-76
~ ~ - ~ -
-wn;m 35 m - - - ~ ' - -- '
1.15P,, membrane tension occurs (25 psi) in the hoop direction, approximately half way up the dome as shown in Figure 4-22. For purposes of calculating conservative values of tendon stress and tensile liner strains, the balance of the accident pressure, i.e., 1.5P a -
1.15Pa , was assumed to be resisted solely by the l tendon network.
At 1.15P,, total tendon stresses (resulting frca prestress increase added to effective prestress) and liner compressive strains up to liner yield were calculated. The liner strains at assumed cracking l are shown in Figures 4-23 and 4-24 as solid lines. The maximum tendon stress at 1.15Pa was calculated to be 137 ksi.
The tendon stresses and liner tensile strains were calculated for l 1.5Pa - 1.15P,. The effect of accident temperature on the liner is to produce compressive strains, whose magnitude depends on the degree of restraint provided by the concrete. For purposes of calculating maximum tensile liner strains, the restraint of the cracked concrete was assumed to be zero and, hence, no compressive l liner strains occur beyond the 1.15Pa pressure stage. The liner strains beyond 1.15Pa were, therefore, tensile and their values l are shown in Figures 4-23 and 4-24 as solid-dot lines. The total liner strain results are shown in Figures 4-23 and 4-24 as dashed lines and the maximum tensile value is 0.000178.
('~
The maximum tendon stress increase beyond 1.15Pa was calculated as 27 ksi. When this was added to the stresses up to 1.15P as a total tendon stress of 164 ksi resulted, which is less than 0.72 x 240 = 173 ksi.
- b. Maximum Predicted Compressive Liner Strains The maximum compressive liner strain occurs when the pressure is at its minimum; therefore, the load condition corresponding to D + F + Pa+T awas investigated. For this condition the concrete did not undergo any through thickness cracking as shown l in Figure 4-25. Liner compressive strains up to 2750F (max Ta) were calculated and added to those due D + F + Pa. The results l are presented in Figurc 4-26 with a maximum compressive strain of 0.002267, which is less than the 0.005 allowable per the ASME
- j. Code.
Using FSAR shear criteria, Figure 4-27 shows that the available shear capacity exceeds the required capacity.
L.
4-6 Revised: 12-10-76 wp. w - ,, ,m -r - - -
._. _ . . . _ . . . _ _ - --mm _. .
~
g .
H
('
4.5-
SUMMARY
AND CONCLUSION The.results of this section indicate that the structural recponse of the delaminated structure to c.itical load combinations would, in general, be l satisfactory. This is due, in part, to considering the actual in place concrete and steel strengths. The in place concrete strength as justified by test ' cylinder results is f' = 6000 psi (see Section 3.3.1) . The minimum l yield strength of meridional and hoop reinforcement based on mill test results is f = 45,000 psi.
However, two aspects of the analyses deserve further comment: l
- 1. the evaluation of the compressive stresses on the inside face of the dome near the ring girder for the normal winter operation load combination (see Section 4.4.2).
and
- 2. the evaluation of dome strength during the LOCA load combination (see Section 4.4.4).
Considering item 1 above, the compression stress due to the normal vinter
- operating condition has been presented in Figure 4.8 with a maximum stress of 3613 psi. The calculated vrlue is acceptable within engineering accuracy when compared to a working stress allowable of 3600 psi. When the
(". compression steel on the inside face of the dome near the ring g!rder is considered and the liner is neglected, the peak concrete compreeston stress is reduced to 3488 psi and the calculated compression steel stress is 25,733 psi (see Figure 4-8). The ASME Code allowable for compression steel ,
for the normal winter operation load combination is 0.67f or 30,000 psi. l 7 l Referring to the second item above concerning dome strength during the LOCA load combination, the analyses presented in Section 4.4.4 indicate the delaminated structure is *heoretically acceptable if evaluated according to the current criteria of the ASME Code. However, reinforcing steel was included in the repaired dome to control and distribute cracking. This reinforcement enhances the strength of the repaired structure and is discussed in Section 5.0. )
I l
r 1
(_ l 7 Revised: 12-10-76
.my _,, .n m---n - - - - - - - - - - - - - -~ ---
l The analyses presented in this section support the conclusion that the delaninated structure would be serviceable subsequent to the repair l activities described in Section 5.0.
This conclusion was further supported by the results of the investigation of the dome and included:
- 1. The measured response of the structure to a 14.6% detensioning was acceptable when compared with predicted response.
- 2. A series of cores drilled into the structure indicated the absence of crushed concrete. There was lower level cracking, but the concrete is sound parallel to the plane of the dome.
- 3. The dome prestress as measured by lift-off tests was greater than predicted, indicating that the delaminated structure was stiffer than assumed and the prestressing forces were not reduced by the delaminations.
x
(.
4-8 Revised: .12-10-76 m# .. . - -
5.0 CORRECTIVE ACTION
5.1 INTRODUCTION
Based on.the analyses and investigations presented in the previous sectiona,.it was concluded that the CR3 dome would be satisfactorily repaired subsequent to the actions described in this section.
The delaminated cap was safely removed; meridional, hoop, and radial l reinforcement provided and a new cap placed. The integration of the new reinforcement with the lower prestressed structure was accomplished to l control and distribute cracking associated with the LOCA load combination.
.5.2 REPAIR METHOD The repair. sequence was as illustrated in Figure 5-25 as described by the following:
- 1. Instrumentation installed and monitored
- 2. -18 tendons detensioned
- 3. Holes drilled (Figures 5-1 and 5-2) into the lower concrete
- 4. Delaminated cap removed i
- 5. Inspection of 24" structure
- 6. Lower leve? cracks grouted with epoxy
- 7. New reinforcement placed
- 8. New cap poured and cured
- 9. 18 tendons partially re-tensioned
- 10. Structural Integrity Test conducted
- 11. Dome surfacing 5.2.1 Instrumentation
- a. Instrumentation for Detensioning and Drilling Grout Holes Instrument stations were established in the dome on two (2) l orthogonal axes at distances of 15, 30 and 45 feet from the apex (See Figures 5-3 thru S-10). The data recovered at these stations consisted of the following: l
- 1. The hoop and meridional concrete strain changes near.the upper and lower surfaces of the delaminated top portion of the dome using Ailtech Concrete Embedment Gages with a 4" gage length.
5-1 Revised: 12-10-76
. c ; y. - - - . .- . - - .
6
.~. .
. l l
'( l
- 2. The hoop and meridional concrete strain changes near the upper surface of the lower portion of the dome using Ailtech Concrete Embedment Gages with a 4" gage length.
- 3. The hoop and meridional steel strain changes on the liner using SR4 three element electrical resistance strain gages attached to the inside surface of liner.
- 4. The change in width of gap between the top and lower levels i of the dome using linear potentiometers (infinite resolution type).
I
- 5. The change in elevation of the top surface measured by survey techniques with 1 0 2 inches accuracy.
- 6. The vertical displacement .e the liner, inside surface, measured at three (3) azinuths and at radial distances of 0, 29, 49 and 56 feet from the apex using extensometers.
- 7. The radial displacement of the liner measured at the 49 and 56 foot radii with the extensometers. ,
l
- 8. The movement of the top of the liner at the apex monitored l by survey techniques using a stainless steel pin attached to I the liner. This measurement was used to correlate the inside l
and outside vertical movements.
- 9. The air temperature at three (3) locations; outside the dome, in the gap and inside the dome. i In addition to these instruments which were added for the repair program, instrumentation previously installed for the SIT test was also monitored. The data recovered consisted of:
The reinforcing bar strain changes on the outside face of
, the dome at three (3) azimuths at approximately 50 feet from the arc =. The gages are SR4 linear strain gages.
The reinforcing bar strain changes on the outside face of the containment cylinder wall at three (3) azimuchs and at seven (7) elevations were also monitored with SR4 gages. l Instrumentation readings except for those recovered by survey techniques were recorded using a Vidar Model Autodata 8 data acquisition system. "The data was printed out at least once every hour.. The survey readings mentioned previously were taken a't least twice a day.
I
\_
N 5-2 Revised: 12-10-76 i
.y._; n .
l
- b. Instrumentation for Other Operations and Repairs During operations prior to SIT, the apex displacement was measured l as well as the_ concrete and steel strains using existing gages.
- c. Instrumentation for SIT The Reactor Building's structural response to the SIT was monitored l by utilizing the existing instrumentation described in the FSAR and supplementary instrumentation installed within the repaired dome (see Appendix J for communications with the NRC).
Instrumentation within the repaired dome consisted of stations l established in the dome on two (2) orthogonal axes at distances of approximately 15, 30 and 45 feet from the apex (see Figures 5-3 and 5-11). The data recovered at these stations consisted of the l following:
- 1. The hoop and meridional steel strain changes on the liner using SR4 three-element electrical resistance strain gages attached to the inside surface of the liner.
- 2. The hoop and meridional _ concrete strain changes near the upper surface of the lower portion of the dome using Ailtech
,. Concrate Embedment Gages with a 4-inch gage length.
.Y
- 3. The hoop and meridional reinforcing bar strain changes of both layers of steel within the new concrete cap. These measurements were obtained from SR4 linear strain gages l attached to #4, Grade 60 sister bars.
- 4. The reinforcing bar strain changes of the #6, Grade 60 radial reinforcement. The gages are SR4 linear strain gages.
Gross structural deformations were measured by extensometers l attached to the steel liner plate at the locations identified on Figure 5-9. This instrumentation provided the following data: 1 <
The vertical displacement of the liner, inside surface,
. measured at three (3) azimuths and at radial distances of ;
approximately 0, 29, 49 and 56 feet from the apex. I i
The radial displacement of the liner measured at approximately the 49 and 56 foot radii.
5.2.2 Dome Detensioning A nymmetrical group of 18 tendons (see Ta,ble 5-1) was detensioned l and the effects on the structure' vere studied. The results of this load reduction is reported in Section 3.
9 5-3 Revised: 12-10-76 wo my,pe -r -
I
('
5.2.3 Drilling Radial Holes The presence of lower level cracks was established by initial drilling (see Section 3.1). While the delaminated cap was still in place and provided a smooth, regular work surface, approximately 1850 radial holes one (1) inch in diameter were drilled into the dome. These holes were located as shown in Figures 5-1 and 5-2. Figures 5-12 and 5-13 illustrate the location of the holes relative to the tendons.
These holes served three purposes. They provided:
- 1. a means for inspecting for lower level cracking.
- 2. grouting and venting holes (see Section 5.2.6).
- 3. holes for placing radial reinforcement (see Section 5.2.7).
5.2.4 Delaminated Cap Removal After the holes were drilled, the delaminated cap was removed to the extent defined in Figure 5-14. Removal work began at the apex in the following sequence, pours Q, P, N, M, L. K, and J, and continued toward the ring girder until concrete above the main delamination was removed and lower concrete exposed.
, Reinforcing steel at the lower edge of the dome (Pours J and H) was
(, cut at the location of the construction joint bewteen pours J and K.
After the loose concrete from pour J had been removed to sound concrete, the final surface concrete was terraced in steps as shown on Figure 5-15.
5.2.5 Inspection of the 24" Structure In addition to the investigations described in Section 3.1.3, the upper surface of the 24" structure was visually in,spected'and boroscope inspection logs were made for the radial holes. The findings of the inspection were:
- 1. There was no evidence of crushed concrete or radial cracks.
- 2. The surface of the 24" structure between adjacent conduits runs generally from near the top of the conduit at lower elevation to near the mid-plane of the conduit at higher elevation. There were a number of localized delaminations forming small lenses (layers) of' concrete usually adjacent to the high side of the conduits. Loose lenses of concrete were carefully removed. l
- 3. Exposed straps attached to the conduits appeared to be in good condition. There was no evidence of any movement of the conduit relative to the surrounding concrete, including those conduits containing detensioned tensons, or of cracks, crushing, or high stresses at the tendon concrete interface.
5-4 Revised: 12-10-76 mmm ;,w a .y - .-
\
The conclusion from these observations was that the 24" structure was in good condition and capable of performing in accordance with the design requirements. Appendix E contains photographs illustrating the condition of the structure during delaminated cap removal. l 5.2.6 Lower Level crack Grouting The presence of lower level cracks was established by core drilling (see Sections 3.1 and 5.2.3). The cracks were grouted with epoxy through the holes described in Section 5.2.3. Figure 5-16 illustrates the device (packer) used during grouting to isolate one level of cracks from other levels in a given hole. A packer was placed in each hole and epoxy grout applied under pressure until grout appeared in adjacent holes or when flow of grout ceased. See Supplement 2, Attachment 2 for additional information.
5.2.7 New Reinforcement To enhance the tensile capacity of the structure to resist the LOCA load combination and control cracking of the concrete in tension, non-prestressed meridional and hoop reinforcement was provided. This l reinforcement is sufficient to resist, at 0.9f , the y membrane tensile forces shown in Figure 4-20. This reinforcement is as shown in l Figures 5-17 through 5-19 and consists of deformed bars conforming to ASTM 615-68. Grade 60. Figures 5-20 and 5-21 compare the area of f'- reinforcement provided versus that required. See Appendix H for cadweld requirements.
The principal radial reinforcement for the repaired structure is a #6 deformed bar and is illustrated in Figure 5-19. This reinforcement was installed in each radial hole with Masterflow 814 cement grout.
Approximately 1850 #6 bars were provided.
The test program outlined in Supplement 1, pages S-ll, S-12, S-13, was expanded to include testing of eleven (11) #6 radial reinforcement test specimens. Test specimens consisted of #6 Grade 60 deformed bar installed in 1"4 hole,15 inches deep, and grouted with Masterflow 814 cement grout. Specimens were tested in concrete blocks with an in-place compressive strength of 2400 to 5000 psi. Grout cubes tested at 4,265 to 7,215 psi compressive strength.
The first observed distress in the test specimens was the development of a shallow secondary failure cone of concrete as illustrated in Figure 5-26. This " secondary failure cone" condition occurred at an average load of 31.5 kips (3.3 kips above the guaranteed minimum yield strength of #6, Grade 60 bar). Failure occurred at an average load of 40.1 kips due to rupture of the #6 bar at welds for testing apparatus attachments. ,
l l
l i
r .
t l
5-5 Revised: 12-10-76
"}e s ,
~ -
l '
The radial tension force for which the #6 bar was designed is a l function of the level of membrane stress in the new concrete.
Figure 5-22 indicates the level of radial stresses in the 24" sttucture, which are all compressive. Future tensile stresses are a function cf the transfer of stresses between the new and old concretes due to creep and recensioning. Figure 5-23 indicates the final predicted distribution of radial stresses at the end of plant life.
The #6 radial bar was designed to resist, at 0.5f y , the tensile force l corresponding to the radial stress of 21.4 psi._ The consideration of radial tension in combination with radial shear is discussed in Appendix 1.
5.2.8 New Cap The concrete materials for the new cap satisfied the requirements of FSAR Section 5.2.2.1. Figure 5-24 indicates the pour sequence used and Figures 5-14 and 5-15 illustrate a typical section. The new material ,
was " wet" cured for fourteen (14) days to minimize creep and shrinkage. '
?
(
l I
. l l
l 5-Sa Revised: 12-10-76
. amwgw??w nwn ~ - - .- - - --- ~- ---
5.2.9' -Dome Retensioning The 18 tendons which were detensioned (see Section 3.1.3) constituted l 14.6 percent'of the tendons in the dome. The average prestress force per tendon computed on the basis of the lift-off data of Table 3-3B was 1491 kips. The design value used in the calculations was 1370 kips. Therefore, the actual prestress forces in the delaminated structure were 9% higher than those used in the design. The force per tendon in the 18 tendons required to return the structure to the design condition wac 646 kips per tendon. l The 18 tendons were retensioned to that load using the reverse of the detensioning sequence given in Table 5-1 af ter the new concrete had been " wet" cured for seven (7) days and had achieved a minimum compressive strength of 4000 psi.
5.2.10 Structural Integrity Test (SIT)
The SIT provided evidence of the adequacy of the repaired structure.
The forces on the structure during the SIT consisted of dead load, pr tressing and 1.15 times design pressure (P,) along with changes in environmental conditions which occurred during testing. Measurenents i and observations recorded during the test sere evaluated and compared with predictions of the expected structural behavior (see Supplement 2, Attachment-1). The repaired dome was well behaved and the SIT provides C -- - quantitative justification of the repair approach utilized.
- 5.2.11 Dome Surfacing A silicone urethane waterproofing membrane will be applied to the l dome surface.
i a
g n
5-6 Revised: 12-10-76 m ws:rrw;wnew.w w - - n~- ~ ~~ ~-- -- - -- -'
TABLE 5-1 DOME DETENSIONING SEQUENCE Sequence Top Middle Bottom
- 1. D101 D241 D301
- 2. D141 D201 D341
- 3. _D117 D225 D317
- 4. D125 D217 D325 5.- .D109 D233 D309
- 6. D133 D209 D333 s
J
~t ,
5-7_ Revised: 9-22-7o n w m, . , ; +wdynwa - sverr-:- ~- ---- v- - -- - .--- -
e 6.0 QUALITY ASSURANCE The Quality Assurance Program as described in FSAR Section 1.7 was implemented during all activities for accomplishing the' dome repair. i This Quality Program was in effect during the entire Crystal River #3 construction phase.
In accordance with that program, approved written procedures were in effect, and were enforced by appropriate quality control methods.
Records of activities and audit of those activities will be maintained.
The Program was supplemented by additional technically qualified l Engineering and Quality Program personnel, with stop-work authority, ,
who were in attendance during all work on the dome. l
(
6-1 Revised: 12-10-76 mrmngm-m.--n-m:- -v~~r - === m e r^-- -- "-
-- - -' - ~ ' ~
e
k
7.0 REFERENCES
- 1. Thomas T.C. Esu, 'Tkthematical Analysis of Shrinkage Stresses in a Model of Hardened Concrete", ACI Journal March 1963, p. 371-390.
- 2. Hubert Woods, " Durability of Concrete Construction", American Concrete Institute Monograph No. 4, 1968.
- 3. American Society for Testing and Material, " Significance of Tests and Properties of Concrete and Concrete Making Materials," ASTM Special Technical Publication No. 169-A, 1966.
- 4. Israel Rosenthal and Joseph Glucklich, " Strength of Plain Concrete Under Biaxial Stress," ACI Journal, November 1970, p. 903-914.
- 5. G. N. Savin, " Stress Concentration Around Holes," Pergamon Press, New York, 1961.
- 6. R. E. Peterson, " Stress Concentration Factors," John Wiley & Sons, New York, 1974.
- 7. William Griffel, " Handbook of Formulas for Stress and Strain,"
Frederick Ungar Publishing Company, New York, 1966.
- 8. Helmut Kupfer, Hubert K. Hilsdorf and Hubert Rusch, " Behavior of e' Concrete Under Biaxial Stresses," ACI Journal, August 1969,
\. p. 656-666.
- 9. Oral Buyukozturk, Arthur H. Nilson, and Floyd O. Slate, " Stress-Strain Response and Fracture of a Concrete Model in Biaxial Loading," ACI Journal, August 1971, p. 590-599.
- 10. Tony C. Y. Liu, Arthur H. Nilson, and Floyd O. Slate, " Stress-Strain Response and Fracture of Concrete in Unixial and Biaxial Compression," ACI Journal, May 1972, p. 291-295.
, 11. S. Timoshenko and J. N. Goodier, " Theory of Elasticity," Second Edition, McGraw Hill Book Company, New York, 1951, p. 58-66.
- 12. Three' Mile Island Nuclear Station Unit 2, PSAR Docket No. 50-320, Appendix SJ.
- 13. Adam M. Neville, " Hardened Concrete-Physical and Mechanical Aspects," (2. Nature of Strength of Concrete), ACI & University of Iowa, ACI Monograph #6,1971.
- 14. Stress Concentration in Prestressed Concrete, Group H Paper 47, From: Proceedings, Conference on Prestressed Concrete Pressure Vessels, Church House, Westminister; S.W.I. 13-17, March 1967, The Institution of Civil Engineers, London 1968.
- 15. Turkey Point Unit 3 Containment Dome, Delamination of the Dome i- Concrete During Post-Tensioning, Docket 50-250.
7-1 Revised: 12-10-76 mamu,v :.v~ rym e m - -
m- - . - - ~ - -- -- -- - --
- , t oe#
y e inQ f.IM /mer ' '
f
.I
,y A k
f fb e
e
, fI. -2&l>Y "
e ,(,
._ - 4 - .__ ,
ev'i .,
, 1 \ ests-
. 2 -
1 l '
Shr fraer if derchet e$
\ _
\ .
/
I
,Q ,,
- LL./72$dh*
A e b N M#.*Ere \
N 3.'s ' / 30' o c a c,gn; KF AAv.re FAgg
'g, i 12 .tifo" I s. }
n -
R / _
2:o"
~ R . u r c a S u i s. n v g g ,_ ,; ,,y ,,, ,,y_ ,
- lG U.h'E .0-l We yn , , ,._
. _ . . . . . .~
(
se@\ o p d) [
.og #j\ 5't 1
, [ ~8S?lo" a o, .
o+.\ o s y \ \ .
,g p 6' \ P l g
L 4*\ ll \
\
\
T G
\H a
- .*8(SHEAR BARS)
/f E +/ -"'M_h .. _
e16 -
D
% f fg*' ,of\~
c WG,3 3 S
PW .,gx Ah
!^OI SHEAR BARS Aa( (512.E %RtES') _
'I
.A NON-PRasTRassso Reissoacemasr l
Fic,ues 2-2 l Revised: 8-10-76 i
- - - , - - - - , , - . , . . - , , ~ - - - . . .
l
,,,n.-
o*-scoct N
n G1 (2257 45* H4-(s.7' 315* 11
% p (g. A.ss%1 (S-t4..y ,,?<> /
g.S (3-2 6 g ,25) pso
( 4. 4 4, 2 )
^
17
\
X \
a6
? 13 o U (7 21 74) v g
5 .
k7 (310'@
/ Js (g3(%
s-\t-'*#
N 13 5* /y,(3-e.74.) 125*
G2 ( ? -15
iso
- Doms CONC RE TE Pl. A CEAdsN T - Pl. A N f/ G U)?E ?- S
. 5. x .;.
O 4_-CONTAJ NMENT VESSEL m P <
g N K i J
R3 H G
I /F -
,- N, g
Dome CONGRETE P&4cEasur 5ecriou Frauas 2-4
, , _ , . - . .. ~ . . - .
b
'- V v(SEE v vv DETAll A"o N)MVW
^
t\N\N\
/
. J \ N W '"
^
v'\/\Mv a s ;
n^^^^g a
iu l Ok PLAN .
\
f CONDUlT STRAPPING g DETA) L F W 0 0/ AA )
\
CINDUN HIFLEY STRAPING DETAIL %" F/s uRE 2 - 5 0
{.-'
,- k'_4 m
SPACED APPROX.
I L (," @ W .P.
yg SPACED d i APPRO't. l'-G"
@ W.P.
- / W. P.
CIRCULAR SUPPORT (TYP) STlFFEWER L SEE DETAIL C* l'd A'% STEEL 4pA O BAR (TYP.)
?oy, PLAN
[
i SUPPORT BAR) i CONDUlT- Li NER SUPPORT DETA)L
\
- TENDON b '
l CON DUlT I" # BA R h
2 f ADI R AL STlFF. L d# $
A l
_ d. .. ..
LIN ER PL l
Suppo aT D ETA \ L Fie unt 2-6
. . n. - . .. l l
l N :
\
/
~
Y
' ece 99 o c>
- e$ o 94o O O D\O\
I\ \\\ \ ~.. \ A I A I A A A 1 I A A A A A N A A A /'# /\/ / / / // /N ,
w i \\ \\ \ \ \AAAIA A A A I I A A A A A A A A API /\1 / / / / //\ 1 A \ \\ \\\ \ / A A A A A A A X I 1 A A A A A A A A Aff.A / / // / //A l A \ \ \ \ \\ \AAAAAAAAA1II I A A A A A A A/\/1 A / / / / / / / A l A\ \ 'i \\ \ \ A A A A A A A I I A I I I A A A A A A A Al\Ar A / / / / / / IA I IT \ \ \ i \\ \ A I A A A A A A I I A A A A A A A A A A A AA A A / / / // / / /A
/T \ \ \ \ \\\ AXAAAAAAIIAAAA gAAAAAAA/AI // / /// A
/\ \\\\ \\\AAAA%AAAAIIA2. A_A .A AAAAAAAA/A/// / / / / /\
A\ \ \ \ \ \ \ A A A A A A AA A A A A A A A A A A A A A A A A A A AA AI / /// // A A \ \ \ \ \ \ \ A A A AAA AAAA A A A A AA A A A AMA A A MA./A/ //////A A\\ \ \ \ \ \ A A A A/bM1\AAAAAA Af\A A A AffiA AAAAA1V/\/ / / // //A
&\\\\ \ \ \ A A A AA/V\AAAl\AAA A A/\A A A APl\ A A AAA A/Vi/V / / ////1 1\\ \ \ \ \ \ A A A/LM/TA A AT1\/VVM A AfAM APA A A AA A/VT.f1V / /////!
l\ \ \ \\\ iA A A AA /M /b% A1V\f/ LV_ A A/V._.NA 7 A A A AA AP/\Pl\/ sf / ////Il 1 \\\\\ \ A A A A/b1/XA AI_AA/V \/vlA A Alb1AA A A AA A AP/\/\/V_/V / / / / / 1 I \ \ \ \ \ A AAAA/\AAA AAr>V VV \AA h/L MA A A A A A A/V\f1N/\PP./ / / / / t I \ \ \ \A AA/bNT/b%AAA1VVVV\/V\/A X/\A A A A/ MAP /\Af\/VV./V / / / 4
\ \ AA A AP/bb\f(A AAP//VV VVV\n \nAA A A A A A?>'JVV\I\/\ffV / / 1
\ \A A A A/VX/b%/\A/\/\/VVV VV'/V L1Ali1\A AA A A/\AA A/\AIbmI/V / ,
\AAAAAImA> AA A AIVV VVVVYVA A A A A A A PA A AME A A/VVV j A A A AA/\/\/\A AliA/\A/VX/\/T/b1W A A AA A A A/\ A A A AA A/VVV \ <
/ A/ 3'A A Al\A A A/\1U\/V71VV V .(VX A A A A A A A A A A A A A AAIV V1 \
/ / A/V\AANIMi\/VV'/\AA A A A AAA A/M AAA.A./\/V \ \ \ l 4 / / / IVb bm_^ ^I .r/L'/ VTP.1\/VVV V 'i(VXA % A A AA.js A A A Af\ AIVV V \ \ \ r n / /// /L'/\AA A/\A1VM'/iAMP/V VVm PAA A A /\A A A A PA A AIb V1 \ \ . s f 1 / / / / / /bViAMaNv/\/VV\/\A/VVVT/V\AAA A AAA A A A A AAP/VT \\ \\ \ l I / / / V \ATA/Vv '.IV vbv/V V Vu\rAIV\ AA A A A A A A r A/VV1 i i \ \ v i
/ // / /\AAfl\/VV/\A/\Alb'IVA/\/V\AAAA A A A A A Afl\Al\ \\ \ \ \ il V/// / / / Alb".1\/\/VV'.1\/\/VL'\/VX/\ ;\ AIV\/\/\/\ A A A A A A APl\/\ \ \ \ \ \ - 7 I V//// / / A/\.AM/T.IVTT.1\/VbT1\/\ A/\Al\/\_A/\/\A A A A A A A APl\ \ \ \\\\V I
\//// / / / AT/V\/ViPEPdVV\/\A /bMAl\A/\AA A A A A AAf\/\ \ \ \\\\V
\1 / // / / / r/V VI\/b'./VVVV\ A A A/\A/\/V\M_AA A A A AAl\/\ \ \ \ \ \\ V
\1 / / / / / I /V V IXAlbVM'VVTA A A>(% A/V i\/\AA A A /\fl\A \ \ \ \ '. \ \/
\ / / I / / / / / V VVV\/V V VV \AA A/MM'VVl\AAA A APl\A \ \ \\\ \ V
\/ / //I// / ~/ VVVVV V '/\/\/bMA/t'/\/1/VVVi A AA/VX \ \ \ \\ \ d h 1/ / I I// I / \/bVVVVVT.A/\/\.A/MA/V\Al\/b31\AAAAIX \ \ \ \ \\ \/
\/ / I /I // IVVV\ VVT./\ A AI\A APJ\.A /V\/VifsA/\Al\ \ \ \\ \ \V V///// / IVVV VVV\A_ XAAA .1A/XAA/\./\A/VV\A \ \\\\ \ V s
\ y V,/,/ , /,I,/ / ,/,V_V v vXn A~-A /\AI\A_/ULA, A/bt,' n ;\ r,IVv vi Nc , x,i~\
ny o> , n V/
&n~ m~ -
& g* Nm o n,N Q
00ME IE NCON l C C A T / O N ,PL A N
, ?/ G UaR i ?-l y,~ ,.
n i,sa*ll x
l ' [i ll
~
!!L
+ l }Cowourr SPUCES
?; (
g h
72 6
gg .B@ s et\1
,9 e. 5
- \\
, -SC.40 PIPE 5"* NOMIN AL 1Y ct.W WALL THICKNESS o.7.58" 0.D. 5.563 I.D. S.047
,/
.g
-L. /
sg L5=3 MERIDIONAL STIFFENER
'O '
' hA- gD XP i I LBx6 CIRCUMFERENTIAL STIFFENER G si! 'e LStB WHERE BOT. REBAR OCCURS
[: - L. CLOSE TO DISCONTINUITY N :l -
/ 3.i hj g LINER PL- DOME DETAILS
ese e m; em e m -- .==esm age = . ew ee =*=. e e
,sa N
APPENDIX D COMPUTER PROGRAM VERITICATION AND DESCRIPTION 4
e b
I i
g g g 6Ggh
- GM O
_ _ __ , .. . .. - e-= -
- MYgd ab,_ '
,% " . ,9 % W -
. _. i KALNINS-STATIC 1
!'g i ~~,
e KALNINS-STATIC ures a multisegment method of direct numerical integration of boundary value problems and was developed by Arturs Kalnins and published in the Journal of Applied Mechanics, Vol. 31, September 1964. pp. 467-476, and in the Journal of the Acoustical Society of America, Vol. 36, July 1964, pp. 1355-1365. 'The progran calculates elastic deflections and stresses in a thin-walled, axisynnetric ' shall when subjected to any arbitrary surface, edge and/or ring loads. The solution is based on the linear theory of elasticity and takes into consideration bending as well as membrane action of the shell in response to applied load. Results are in terms of resultant forces and couples with stresses calculated by assuming a linear distribution through the thickness.
This program has been widely used for thin shell analysis since its release 3
- to the public domain in 1968. The program is being executed on Gilbert Associates, Inc. , Reading, Pa. , IBM 370/155 computer under INI operating system 0/S 21.8 MVT with HASP 3.2.
e I
D-1
~
- + -- - -. - - - ...
- + "OYY9/
- N' *b #" P T F?O M veg g 4M., y- ,4,,q s . _ ,
SAP IV
( r5
'-- SAP IV is a large scale linear, three-dimensional general purpose structural analysis program having static and dynamic capabilities. SAP IV has an extensive finite element library. Nodal points can have six degrees of freedom. No restriction is made on the number of nodal points or finite elements used. Temperature, hydrostatic, inertia and surfaca loadings may be used. General orthotropic material properties can be modeled.
SAP IV wiss developed by the University of California at Berkeley. The program has'been widely used for structural analysis since its release to the public domain in 1973. The program has been verified by comparison to published results.
It is being executed at United Computing Systems, Inc., Kansas City, Missouri on multiple CDC Main Frames (i.e., 6600, CYBER 7418, CYBER 175) under APEX Release SL7.
~
'A t
e O
s
\
e
~
D-2 e -
4,.' M 7 N[__ -[ p.'p7 g
,9 * * * ' ' '""hW** -**"'t"-*h"'*"
-. ~
~~
. . . . .C._ . -
ARTURs KALNINS
( t}' @ Ja , 6 , w N.O.3 34N64N N040 SETMLEMEM, PENMSYLVANIA 18095 U.S.A.
m esis e ses.si e May 11, 1976 Dr. F. L. Moreadith Gilbert Associates Inc.
P. O. Box 1498 Reading, Pennsylvania 19603
Dear Dr. Moreadith:
Upon your request, I have completed the running of 18 test cases with the KALNINS Static Computer Program for the analysis of axisymmetric shells.
The test cases were designed to test the following features of the program:
- 1. Geometry of a shall (standard surfaces)
(' , 2. Elastic foundation s \; 3. Orthotropy of material 4 Variable surface loads
- 5. Variable thickness i
- 6. Edge loads
- 7. Springs
- 8. Branches and box branches
- 9. Ring loads
- 10. Layers with different material
- 11. Dead weight loads
- 12. Torsion j
- 13. Subcase addition 1 I
In all cases, the results given by the program agreed with independently cal- '
culated values which were c.htained either from published results or from a global equilibrium requiremaat. All cases were run first on Lehigh Universi.ty's.
CDC 6400 computer and then rarun on Gilbert Associates IEM 370/155 computer.
In all cases, there was agree;3ent between the two (2) outputs.
I believe that these test cases provide a satisfactory verification of che validity of the KSHEL1 computer program.
Sincerely yours,
. l l Q yk {//! 1 '
Arturs Kalnins GFD:AK: fig D-3
~~
'g*. Ya & +'t
~l
(
APPENDIX E PHOTC';RAPHS OF EXISTING STRUCTURE 1
1 1
l l
l
=
l 1
1 l
l l
s_
Revised: 9-22-76 1
1
~~~ *""#~~' = ~
'Jhyh +, hh'g:L .h ,y. . 'gr -
APPENDIZ E f
PHOTOGRAPHS OF EKISTING STRUCTURE phoro NO. 1 -
Surface of Concrete Illustrating the Relative Position of Failure Plane and Conduit PHOTO NO. 2 -
Surface of Concrete Illustrating the Position of Failure Plane a.ad Lens on Surface I
?
s.
l E-1 Revised: 9-22-76 l
. mn . _. - . ,,ww o n-- ~. --- ~ ~ ~ ~ ' ~ ~ ~ ~
- ~~
.yg ~ .srit 1 tif-I d-
.. Av w 0 a
. ?' ne m, .
ts?
b' M
c.. Q 2 A y,n .k^iG Q, ! ,{ }A 3 3 h'g
% 2%.d L* ~ e,
%c. ,,; . u %,74; ; ki!!n: k ;
- f
- g'e f.J;-l. 3gg.. .
.wiyh;t T.pV;e,aq)gQQ;. ?A%aiw s, ,M H'? Nke LW4 w: WdVyF v.
g.;
ep yfwn%,,&n >
= ,:np A y Q3y yg7; e w s4,:M$;y;y %6
%Qp-~g%g
.p , Q, ; .
'iM:" 's.
c;GSC W ; @ -..
- ' ' .'s '% W9 m:s;m:@
( ** f 4 !&dNA. jeCxg.:
l PHOTO N0. 1 l
\
l l
l 1
a Y
l -
+
.N
, ;- +
~
.A..'
~
3.. .
,;-c ,.
gy -- -
- . * .- ,. . g
.. . .Q ,, .w g ..
i
. .. .i .
, y .
, . w> . . . . .,,.. ,
s, >
' . ' +. . F..e, ;f .g.. *; ')sh., $
.. jg ..d
{' I '
y %g)qf .
[:%W
- A. N .; . . , --
W :-
}
. ; ; - n ..
_A
.n,$
, , pc,w : x.@,,. (_
n -
1
^
yn :4? h# n
., .arJy.nun:Z:e [+.
G n -;
I n
. :?q4 7p%&M, stiw 84lf'ff i(
%.dr - -~g& ~":4;%y-2+
ed$$y*l~Q;&[W l
PHOTO N0. 2 1
I l
E-2 Revised: 9-22-76
~ , . ~ _ . . ~ . - . ..
-- APPENDIX E (Cont'd.)
PHOTO NO. 3 -
Loose Concrete Inns Extending to the Top of the Tendon Duct Phoro NO. 4 -
Tight Concrete at the Top of a Tendon Duct (at Different Location than Photo No. 3) i l
l l
1 E-3 l Revised: 9-22-76 l 1
l 4dermm -_ _ _n _ we -=s.. ' + - - < - - - - - -.--- -.,, a
l l
t 1
- m. < -p - s s ;
...,._,..i.:-2
-s. .a , gy;- .
- r. '.. v 4%..~k. -
, k., g f,, 4. - .' n *
.@ . .. ;r
- n , 1 et
,, ..'s .
g: ;
+p _;. .
.2 ;
a
.s .
_%.T< i ! A. l -
,3
. ,tf.
4 ; = .r -
j
- r. ,*;-
~
9,94 '
y 2 - --
r>
s .# y
, .s A - R.s*
s .
r
- %v :,:
0.,.
' +
% ..,s .,
w
_). s-
~
- r. n g 4 . f-
, f f' ;
' g. '. f - ' n.
.-+s
+., 3' .
.A,.
@J- , -
. , - .. .o , , - - ..
.v. ; , ,, . - g . - +
g, r.
4 g.). 4 .-.
- 5 1 3 7.. ; , '3
~
.g s,g. . . . , ..
. [* ' ,'fN' , ,, ~ g hs , k
- a. . +; . y
. ; g. .p s.
_. y.
.c py,3
';. . , - 3 , .- ,
. ,t. -
,~~
- - ' 4._u, h 1..
J. .' $. .f . _ . g?')
( i
.h,.f '
[ '
9,'*
l I C>s@ -
' . y ^gy,
. , d . d .2^ 5 ' .1/6.q ' . 9 % . *. __ 42 ' : :
. . , ..: . . :". ., s
\ ..
! dY Qi &,
\~
% s '/ PHOTO N0. 3 4'6' W &hY u
- ..- ' * ' y .- .
' y
- w. .~ +
'.,'. ?
Q'.L,, ,
- - l , .t ' :'
z.
~
-- e
- . ,.s .
e .
. ~
j -. ; -
~
~
.c.
-.+~ .-_
l 4 ,. , 4
_ ., w
-e ,
(;; - ,, y, f
< ? <
i g +' ', %2 ;~ .
j
-y? g -
'4
<, e ,. +.
99* ..f, y,
~
PHOTO N0. 4 E-4 Revised: 9-22-76
.- ~ . s - w ,m - ..,~ _,
, APPENDIX E (Cont'd.)
i PHOTO No. 5 -
Close up of Delamination as it goes across a Tendon Conduit x
9 E-5 1
-er..,_. - . , , . _ _ . , . . _ _ _ _ _ . . _ _ __ _ _ _ _ _ _ _ _
i I
i i
p 4
. 7.s gf Y !!
- . P u_. J e.f 4 .
t,,ag,w:,x;n s 3
.. -l,,.j p' ;1:u ;a : ,
n .:. A w . . . . , .n ; x. .,,8
- . :n , ;,) xv ,
r;a ,.a
' ' y].',* $'(
^
s w +y,.
. A eJ ,.
. ye
..,sg Ci 1
5 p , y;. , h; thbc..gT3y y ,, fb' yi w ,
h;
. . , c .. , ,
43 R. . ,% .;' f, ,
y
- jw* h .-._ yfs t :^
. u. .
. , egy PHOTO NO. 5 C o %g- t
\
9 k 4%D W
eh~
I f 4 i
E-6 Revised: 9-22-76
. - , - - - - - - , _ n. . - - - - - - - . , . - - .
, . , - - , . , . - - ,,-n., , . - - - . _ -._...,.,,..---.-_-_.-._..__.-.,_,_n.,,_
t
. APPENDIX E (Cont'd.)
M NO. 6 -
Botton Surface of Delaminated Cap PHOTO NO. 7 -
Bottom Surface of Delaminated Cap a
E-7 ,
Revised: 9-22-76
.y.,. _ _
-.-..m... .__ . - _ - _ _ . . _ _ _ .,,,, . _ _ _
- w. . we n -
& N,.f'lA s; . ey _ , g w ' ?.:. 'sc' f
i ., . _' +
h- .g v. j.U '.*,th 'y ..N 3 ..
m
- +
, 4
. % 4.' , m,w%~. p.,
av
-g%.,'.'S
-f -
. , 2 # ' 'h~f" ,
- . )
g#.p.3 . 1';1.'c-
.' g .
1
. . : o _-
- . ~: :
r, .
..g m
, pea.m r. - .
., *. . yg . ' ' "' ? . . ;: .:
r ..
L
.e :.
x I.h* ( , ,: . ; .
p j ,
I
,,, s,.o ,s- ,. .;.
f,
~
~'
)
i.iumA hist - .- . - - ~ -: :.. .=
I i
PHOTO NO. 6 i
r ;7 m
, kr 3.. ., 9, * < ,,,
- - ,.p m, ..y,ye.
.r. ,: w
- t' *,
, m yg W
'm .::r e .
. is V A . -... ..
3c a.q : - ..
. f j. ',D' $' $.p. -
- 9 s'; *7
..3
.s E .
[r ..,s p
.. u;,;,; .: s .. ;.-
-;x. o r 9. .>c. ; .; -- y.c .
s s ,
.-..,s.,, ,
g ,
i
.Y ' ' ' . r
. . . -.fo J;.m.% % '
'/, __ :4)g=' : . .l. D Q ;
N'
- l .y}fpg,k ; ;y, , ~ .1
. , q, .
., .d ,
o
.; a y ftQ. ~ f.4pi;d.pf.g. -2.,,
.v 4
-o ;g 4 --..
A
.,;4u .f ..
,; ',ys '+tW W W f, .
.we f: . '_.,
=*
F: )
h{h , . .; N
. j' '
._ . :-[ ~;., f '~
, ~[ N I
l i
l i
l PHOTO NO. 7 l .
E-8 Revised: 9-22-76
..e--,,, .- -
-e~~.---..- ,
, - , - l
5 4
3 3
3
}
(
i APPENDIX F DISCUSSION OF REPAIR METHOD 1
I c f'.
I Revised: 9-22-76
. ~ , - - ._ _ .
**'h58P*-- ,
h= s. .,
I
APPENDIX F DISCUSSION OF REPAIR METHOD
, Hansen, Holley and Biggs Inc. consutrIno una1unnas I.
FOBERT ). HANSEN
- MYLE J. HOLLEY, JK.
- JOHN M. BIGcs Box 88, AllT Branch PO, t~ambridge, Atassachusetts 02139 August 6, 1976 Dr. Fred L. Moreadith Gilbert Associates, Inc.
P.O. Box 1498 Reading, PA 19603 Re: Crystal River #3 Dome Delamination
Dear Dr. Moreadith:
This is written to summarize my present judgments relative to the subject structural problem and proposed corrective measures.
CAUSES OF DELIMINATION The delamination represents a concrete tensile failure in a tri-axial stress field comprised of two substantial compression stress compon-ents and a small tensile stress component. The averaae tensile stress across the failure surface was modest (= 40 psi), but it reflected an
( " active" radial force; i.e., a force unrelieved by local cracking.
There were many sources of lochi stress concentration which, in some combination, could produce local stress states exceeding the concrete strength.
It appears that the concrete used in this structure fractured through i the aggregate when loaded in tension. This suggests the possibility that the concrete is unusually " notch sensitive," that is, subject to tensile fracture in regions of high stress concentration. In the presence of the active radial force, such local fractures would trigger propagating cracks.
The head of such a crack is itself a region of severe stress concentration, in a field of average tensile stress which increases as the crack spreads.
In sumary, a) Geometric discontinuities and reductions in effective section, occasioned by the conduit grid, easily could produce locally con-centrated stress states in excess of the strength of this concrete.
b) The concrete was unusually notch sensitive; that is, likely to fail in a brittle manner at points of local stress concentration.
c) The radihl tension force was essentially unrelieved by the cracking, until the cracked area and crack thickness were large. Indeed'as the cracks spread, thereby mducing the area resisting the radial
, force, the magnitude of tne average radial stress was increased.
F-1 Revised: 9-22-76 gn. _ _ - . - _ . r- , _. . - - --
Appendix F Cont'd.
('
d) Under the impetts of the unrelieved radial tension force, a crack initiated at a point of stress concentration (and subsequently generating its own point of stress concentration) would spread very rapidly.
CONDITION OF REDUCED DOME Thus far exploration of the concrete in the reduced dome has been by a substantial number of core holes driven through the cap into the reduced dome below. Secondary laminar cracks have been found, particularly in regions away from the boundary. Apart from these cracks, the exploratory coring has not disclosed any signs of distress; e.g., crushing or spalling of the concrete. Indeed the fact that the core drilling has not, itself, caused any evidence of spalling may be an indication that it has not pene-trated any zones of unexpectedly high compression stresses. I do not know of any studies of damage due to core drilling into stressed concrete, as a function of stress level. Nevertheless, I would expect such damage to occur as the stress level approaches the cylinder strength of the concrete. The absence of any damage associated with the core drilling, to date, is at least encouraging. It may or may not verify that stresses do not exceed the calculated values, which are well below the cylinder strength.
It is my impression that the lift-off forces (in the 18 tendons checked) imply, on the average, smaller prestress losses than would be pre-dicted by the creep relationship used to predict 40-year effective prestress values. This is further, indirect, evidence that the concrete in the re-duced dome is in generally sound condition. .
Additional verification of the sound condition of the reduced dome must await removal of the old cap. This will permit the first, direct, visual inspection. If the results of that inspection are favorable, repair measures which utilize the existing dome will be justified.
If visual inspection of the dome does not disclose any severe, unex-pected, damage, there would appear to be no necessity to de-tension the tendons. Indeed, the present stress and strain condition of the structure must reflect a history of inelastic, as well as elastic, strains; and the response to de-tensioning (essentially elastic) cannot be expected to re-turn the dome to a zero condition of stress and strain. Thus de-tensioning might cause further cracking, might open cracks which now are small, and might cause separation of conduits from the concrete. These are undesira-ble effects which should be avoided if possible. Unless stress analyses indicate that the tendons must be detensioned (and re-tensioned after a
- new cap has been joined to the dome), or unless visual inspection reveals unexpectedly severe damage, de-tensioning is not recommended.
- REQUIRED DOME BEHAVIOR It is obvious that delamination caused the prestress forces (only
(, slightly reduced) to be resisted by a dome of lesser thickness. That is, F-2 Revised: 9-22-76 ww- - - M * % T _5fp % _g,p
.-y
Appendix F Crnt'd.
L' it caused a significant increase in meridional and hoop membrane stresses.
Since the conduit zone is a larger portion of the thickness of the re-duced zune than of the original dome, the distribution of membrane stresses through the thickness also could have been altered. For example, stresses in the zone between conc.Jits and liner might have been increased by this lat'.er effect as well as ey the reduction in dome thickness, per se. To whatever extent such changes in the distribution and intensities of membrane stresses may have occurred, the accompanying membrane strains could lead to altered shear and bending moment distributions in the discontinuity zone at, and near, the perimeter.
As already noted, there is no evidence, in the form of crushing or spalling, to indicate that the stress changes in the reduced dome have in-cluded the development of regions of very high compressive stress. Moreover, the relatively small tendon stress losses during the 19 months since de-lamination sugge:t that abnomally high concrete stresses have not developed.
This seems to indicate that the conduit zone is sufficiently stiff to accept a substantial portion of the membrane forces.
Under accident conditions, it is clear that the inward radial pressure which can be developed by the tendon grid is substantially in excess of the active outward pressure (1.5 x 55 = 82.5 psi). Preferably the inward and outward pressures will be in balance at a membrane strain which does not imply the possibility of large cracks in the concrete. In verifying that this desired condition will obtain, it appears prudent to assume that the concrete has zero tensile strength.
To achieve the above-described desired behavior under accident load conditions, it is essential that the tendons not experience any unexpected, large, loss of tension. This, in turn, implies that the dome must not undergo any unexpected, large, compression strain. Lif t-off measurements .
have shown prestress losses, to date, to be modest. This, in turn, demon- l strates that no unusually large concrete compressive strains (elastic, and/ !
or inelastic) have developed thus far. For this reason one may have confi- '
dence in analytical predictions of the prestress losses which may occur over '
, the next-40 years. Such analyses indicate that the residual tendon forces will be sufficient to cause inward and outward pressures to balance either at a small net concrete compressive stress, or at a net concrete tensile stress within the capacity of bonded rebar which will be placed in the new cap. Accordingly it appears that uncontrolled concrete cracking will not occur under the accident condition.
In addition to avoidance of uncontrolled membrane cracking, it is essential that the dome possess adequate resistance to the radial shears and bending moments which may develop, in the transition zone, during the accident load condition. Fortunately it appears that this zone suffered very little damage (e.g., secondary cracks) as a result of the delamina-tion. Your analyses appear to indicate that the shear and bending effects are less severe in the accident condition than in the nomal winter operat-ing condition. If the transition zone' doss not experience any " failure" during nomal operation, it should be well able to resist the shears and bending moments associated with the accident condition.
F-3 Revised: 9-22-76 w - e . ~- .. n - --
App;ndix F Crnt'd.
.e
~k
.Let us next consider the behavior of the dome during normal operat-ing conditions. In regions well away from the boundary the new cap will slowly assume some small portion of the compressive membrane forces cur-rentiy being resisted entirely by the reduced dome. Thus the stress state in i.he reduced dome in these. regions of essentially membrane be-havior canLonly improve with time. Because of the absence of substantial shear force in these regions, the secondary (laminar) cracks are not a threat to the membrane compressive strength of the repaired dome. Even without the addition of radial anchor bars the region should not experi-ence any distress. However, the contemplated radial bars are required to mobilize the membrane compression capacity of the new cap, and they will improve the membrane compressive strength of the concrete in the reduced dome.
In the boundary zone your analyses indicate high meridional compres-sive stress due to the combined effects of membrane force and bending
' moment. This stress, which is the subject of continuing analyses, may be
.slightly in excess of the 0.6fc FSAR limit. If so, th's does not neces-sarily indicate that the dome can not safely withstand normal operating load conditions. First, it should be noted that the meridian compression i stress in question is only 20 percent smaller for the D.L. + prestress
. . condition than for the normal winter operating condition; the D.L. +
prestress condition has existed for 19 months, and has not caused any evidence of distress. Second, the computed high compression stress re-sults from linear analysis. Analyses which account for increased elastic and creep strains with-increased stress should indicate some reduction ir. l the section bending moment, more favorable distribution of the compression stresses on the section, and greater participation of the m'eridional com-pression rebars. All of these effects should lead to a lower value of n computed stress. Third, it should be noted that the load (mainly D.L. +
prestress) is very reliably known for the operating conditions. Under these known loads the section bending moment is well below the ultimate L section capacity. Finally, it should be remembered that the high com-pression stress. vanishes under the accident condition.
Based on the immediately preceding discussion, a) The actual maximum meridional compression stress in the boundary zone probably is less than the computed value, and may be below the FSAR limit of 0.6 f' .
ress
' b) The high computed catastrophic failure comp (e.g. ionunder
, collapse) stressnormal does operating not signify any threat of condi-tions.
PROPOSED flew CAP I am in complete agreement with' the decision to remove, and replace, the old: cap. tiot only will this permit better inspection of the redoed dome, but it will permit the installation of an improved grid of bonded rebars in the cap zone. The quantity of meridional rebar which can be -
l F-4 Revised: 9-22-76 mi frW:TM .= n n_n,, .,
~
n.
Appendix F Cont'd.
(' developed is limited by the existing quantity of these bars outside the cap boundary. However,-the quantity of hoop rebars can be increased.
For the accident loading there is a region over which hoop membrane force exceeds the concrete pre-compression. That is, computed net concrete stress is tensile. The increased hoop rebar quantity will be sufficient to resist these net tensile forces, thus precluding unconstrained growth in crack widths. Finally, Cadweld splices, rather than lap splices, will
,- be used for rebars in the new cap.
It is my understanding that the new cap will be of essentially the same thickness as the old cap. Because the new cap concrete will be placed on the prestressed reduced dome (i.e., with only 15% of the dome tendons de-tensioned), the membrane compression stress in the new cap will be'substantially less than in the old cap (prior to delamination). Re-tensioning of the .15% of tendons that have.been de-tensioned will produce only small membrane compression stress in the new cap, and this will be reduced by shrinkage strain. It must be expected that the structural integrity. test loadirig will produce net tensile stress in the new cap in excess of.its tensile strength; thus cracking will occur. Upon removal of the test loading the membrane stress probably will again be compression, but small. Creep strain, over many years, may be expected to cause a
, modest increase in the cap compression stresses and a corresponding, but
- smaller, decrease in membrane compression stress in the reduced due.
I have reviewed C. Chen's analysis of the maximum (i.e., long-term) i membrane' compression stress that may develop in the new cap. I agree with
- ( his conclusion that the cap compression stress will not exceed 1200 psi; it may be substantially less. This implies that the average tension stress across the cap-to-dome interface will not exceed 21.4 psi, in regions where j the cap is '12" thick. In regions of lesser cap thickness the radial stress will be proportionately smaller. It is my understanding that the radial anchor bars securing the new cap to the reduced dome will be proportioned i for the above dome radial tension stress, at a bar tension of 0.5 f . This
' is reasonable and conservative. Y C' ORE DRILLING, GROUTING, AND RE-DRILLING
- . Most of the secondary cracks are in regions away from the boundary,
! that is, they are in regions of essentially membrane behavior. For this
- reason they may be of little or no significance to the dome strength.
. Nevertheless, I believe the decision to (epoxy) grout all of the secondary cracks is' prudent. Unfortunately this operation together with the radial
, anchor installation will require repeated core drilling. The crack grout-ing operation will fill the radial' holes drilled for that purpose with grout. A second set of holes, essentially co-axial with the grout holes, but of larger diameter, must be drilled to accomodate the anchor bars.
j I 4m not familiar with any research on the effects, if any, of drilling
- small diameter holes into biaxially compressed concrete. However, there has been no evidence of any damage associated with the substantial-number of exploratory holes drilled into the dome thus fo; . Despite the fact
~
.that the'new cap will develop smaller interface radial tension stresses I than were associated with the old cap, the fact of the earlier delamination F-5 Revised: 9-22 . . . .. . . . _ .
- , f s -, -
[h h . . . .y * @ AgP '$ -- - - _ - g -gg g '.t , ,.
wm,w, ,.
.r y , , , , _ . _ _ _ , . , _ _ , ,,__._,_,_.....,.._.._..._,_,,,__-.._...m ,_ __ . . , , ..
Appendix F Cont'd.
( dictates some positive radial connection of the new cap to the reduced dome. Paul Mast's suggestion, that the new cap be tied to the outer conduits, is interesting, but it is not clear that there is a reliable practical means to accomplist. this. Therefore radial ties appear to be necessary, and it is advantageous to extend them deep enough to cross the secondary cracks in the reduced dome. Accordingly the risk that the drilling may cause some damage to the stressed concrete appears to be both small and justified.
SUMMARY
- 1) The decision to utilize the reduced dome, unless visual inspection reveals unexpectedly severe damage, is sound.
- 2) The decision not to de-tension the tendons, beyond the present 15%
de-tensioning, is sound. -
- 3) The decision to cast a new cap of essentially the same thickness as the original cap, and to place adequate, Cadweld-spliced, bonded rebars in the new cap should assure satisfactory behavior of the dome.
- 4) The radial tension stress across the cap-to-dome interfacc nas been conservatively computed.
- 5) The decision to grout the secondary cracks is prudent.
- 6) The use of' deep radial anchor bars, tying cap to dome and bridging the secondary cracks is sound. There is no basis for certainty that the necessary core dr.iling will cause no damage to the compressed concrete. However, the risk of such damage seems small, and is justi-fied in view of the advartages gained by the use of such bars.
- 7) It is believed that continuing analyses may show that the high con-crete meridional compressive stress near the boundary is not excessive.
However, regardless of the results of these analyses, it does not appear that a modest overstress in this zone has any adverse implica-tions for the perfonnance of the containment structure.
I look forward to the opportunity to discuss all the above with you and your colleagues at our next meeting, on August 9,1976.
I Sincerely,
? .. ,,
MJH/jm Myle J. Holley, Jr.
i F-6 Revised: 9-22-76 h'7 , , 7--NNTi ',.. _
_ @N
(
APPENDIX G COMPARISON OF DESIGNS 4
Revised: 9-22-76
~~ -_ _
-.1 _ . _ . . _ . _
4 APPEi._iX C
! Pega 1 of 9 i COMPARISON OF DESIGNS 4
SUMMARY
OF RESULTS i
} STRUCTURE: 36-Inch Dome (FSAR - 5000 psi Concrete)
?
1 Dead Load Norm. Winter
+ Operating Prestress Condition Structural
, Item @ Early @ Early Integrity LOCA s No. Description Plant Life Plant Life Test @ 40 Years a
Design compressive strength of concrete (psi) 5000 5000 5000 5000 Design dome tendon prestress (ksi) 148 148 148 137
-1.56 2 2
- Deflection of dome apex (in' -1.52 +0.75 3 +1.02' .
- . __ -r-m mmm - - _ . - - - . - - - - - - - - - ~ x-- - -
Maximam tensile liner, strain (micro in/in) None None None None I' i Allowabic tensile liner strain 1000 1000 1000 1000 j
(micro in/in) I i
-l Haximum compressive liner strain '
948 1240 510 2057 t
III (micro in/in)
Allowable compressive Ainer strain 5000 5000 5000. 5000 (micro in/in)
' i Pressure at which first concrete cracking No Thru -
No Thru No Thru No Thru 3 i occurs through section thickness (psi) Cracking Cracking Cracking Cracking IV Maximum pressure for transient (psi) 0 0 63 83 ;
% i; C*
d 1 Maximum calculated tensile stress (148 psi) is less than allowable value (+212 psi) .
2
, Referenced to unstressed condition. .
fj 'Due to 1.15 Pa h "Due to 1.5 Pa+Ta l
1 Appendix C Cont'd. Pega 2 cf 9 j SUFetARY OF RESULTS STRUCTURE: 36-Inch Dome (FSAR - 5000 psi Concrete)
Dead Load Norm. Winter LOCA
+ Operating Structural @ 40 Years Item Prestress Condition Integrity No. Description @ Early @ Early Test
- Plant Life Plant Life i.
Maximum concrete membrane compressive stress (psi) 1836 1759 Maximum concrete membrane tensile. ,
V stress (psi) -4691 148 Allowable membrane stress (psi)
- -2250/0 - -2250/0 -2250/0 -2250/+212 *^
- (compression -/ tensile +)
_, _ .___ _ _ :. . ___----.~_..____--=____.- _-__ --- -
i Maximum compressive cmcrete extreme fiber
. stress (psi) - Unc mcked analysis value 2666 3538 ,
i Cracked section investigation value - ,
VI Calculate when actual uncracked is Not Reg'd 3038 ,-
greater than allowable Allowable compressive extreme fiber 3000 3000 3000 3000 st .as (psi) ,
~ ~ ~ ~ ~ " '^' ~ " " ^ ^ ' ~ ~
=x - _ _ . . ._ ,_
ps Maximum concrete tensile extreme fiber None None None None
$ stress in areas without reinforcement (psi) e A VII l Allowable concrete tensile extreme fiber 0 0 0 424 e stress in areas without reinforcement 8
y (psi) i L
os I
No tension occurs; minimum compression indicated.
l '
i
^
l m
.i I
j ArPendix G Cont'd.
P ge 3 cf 9 1
1 SLMIARY OF RESULTS STRUCTURE: 36-Inch Dome (FSAR - 5000 psi concrete)
Dead Load Norm. Winter MCA ur ears Item Pres ress nd tion No. Description @ Early 0 Early Tm Plant Life Plant Life Calculated shear at point of minimum 53 17 15 10 margin (kips /ft)
Allowable shear at point of minimum 55 25 23 23 VIII margin (kips /ft) j Location of minimum shear reinforcement 153.8 158.4 155.6 156.6 margin .
= _ _ _____ _ - -- --- ---- -- ---
_ m e:acesus - - - - -
i Maximum tensile stress in reinforcing Section bar (psi) Not Cracked 2192 =0 25125 IX Allowable ' stress in reinforcing bar 20000 20000 20000 36000
! (psi)
?
. E i
N b
m
(
Pzge 4 of 9 2 Appendix C Cont'd.
I
SUMMARY
OF RESULTS
-1 STRUCTURE: 24-Inch Dome ("In-Place" Concrete Strength - 6000 psi)
, (See Table 2-2 for allowable criteria) '
1 1 Dead Load Norm. Winter
+ Operating Prestress Condition Structural Item ,
@ Early @ Early Integrity LOCA No. Description Plant Life Plant Life Test @ 40 Years l Design compressive strength of concrete 6000 6000 6000 6000 (psi) t 141 141 141 125 Design dome tendon prestress (ksi)
Deflection of dome apex (in) -1. 8 3" -1. 7 9' +0.f0 +2.07 8 -
m _ _ _mm m_ - . , - - _ -
- - - - - - . m _ ---a=a= - - - - - - - - - - - --- - -- - -
Maximum tensile liner, strain (micro in/in) 2 None None 178 II
. i Allowable tensile liner strain 1000 1000 1000 3000 .
(micro in/in) L l
Maximum. compressive liner strain 1100 1560 570 2267 III (micro in/in)
, Allowable compressive liner strain 2000 2000 2000 5000 l (micro in/in)
__ _ _ _ _ _ _ _ w __ __ _ _ _ _ _ - - - _ - - - - - - - - - - _--- -- -
i
- Pressure at which first concrete cracking No Thru No Thru No Thru occurs through section thickness (psi) Cracking Cracking Cracking l IV y Maximum pressure for transient (psi) 0 0 63 83 I
M r
o R ' Referenced to unstressed condition.
2 l"e Due to 1.15 P a 3
h Due to 1.5 Pa+Ta -
Y t
lu Appendix C Cont'd. Pcge 5 of 9
) ,
SUtefARY OF RESULTS STP.UCTURE: 24-Inch Dome ("In-Place" Concrete Strength - 6000 psi)
. _( See Table 2-2 for allowable criteria) 1 Dead Load Norm. Winter LOCA j + Operating Structural @ 40 Years
> i Item Prestress Condition Integrity No. Description @ Early @ Early Test ,
Plant Life Plant Life t
I. Haximum concrete membrane compressive g l stress (psi) g Maximum concrete membrane tensile-V stress (psi) -5851 4452 Allowable membrane stress (psi) l (compression -/ tensile +) - 00/0 -2700/0 -2700/0 -2700/0 ,e
,_ ,y _ _ _ _
? Haximum compressive concrete extreme fiber stress (psi) - Uncracked analysis value 3266 4082 l
Cracked section investigation value -
VI Ca'1culate when actual uncracked is Not Req'd 3613 i nreator than allowable -
Allowable compressive extreme fiber stress (psi) 3600 3600 3600 3600
..__m. _.- - - - - -m -m--- mm e= z -- - ---_=- >
Maximum concrete g n_s*1e extreme fiber 9802 y stress in areas v thout reinforcement None None None
- Uncracked) 4 (psi)
{
VII Allowable concrete tensile extreme fiber ,
j stress in areas without reinforcement 0 0 0 0 y (psi)
M y
i I No tension occurs; minimum compression indicated. -
2 Since calculated values exceed allowables and no reinforcement exists, LOCA resisted as described in report Section 4.4.4.
.l .
1 s
Pegn 6,of 9 i-
~
Appendix G Cont'd.
SU!O!ARY OF RESULTS STRUCTURE: 24-Inch Dome ("In-Place" Concrete Strength - 6000 psi) 5 Dead Load Norm. Winter LOCA Item Pres ress nd io [g ,
No. arl a Description p p Tes" Calculated shear at point of minimum i
margin (kips /ft) 45 55 .6 10 Allowable shear at point of minimum VIII margin (kips /ft) 55 55 7 23 Location of minimum shear reinforcement 151.5 152.00 161.0 margin 160.0
- _ _ . . _ _ - . _ = . _ _ - -. - , _ . .
, = = = = _ m. nag;n - -- --- -- - -- - _ _ _ _ - _ _ = -
i Rebar not Maximum tensile stress in reinforcing 748 7081 available
=0
, bar (psi) over most IX of dome i Allowable stress in reinforcing bar 4
(psi) 20000 20000 20000 36000 i
i (See Table 2-2 for allowable criteria)
!e3.
i i U ih I
i
. n
)
3 Page 7 of 9 Appendix G Cont'd.
j i
SUMMARY
OF RESULTS STRUCTURE: 36-Inch Dome, pr.luding New Reinforced Cap Dead Load Norm. Winter
+ Operating Prestress Condition Structural i Item , @ Early @ Early Integrity LOCA No. Description Plant Life Plant Life Test @ 40 Years Des gn ompressive strength of concret 6000 6000 6000 6000 i
i i
Design dome tendon prestres (ksi) 141 141 141 125 3
Deflection of dome apex (in) -1.87 -1.83' +0.90 .' 3
+2.07 8 5 -
.f Maximum tensile liner. strain (micro in/in) =0 0 0 1781 I. 11 T3 Allowabic tensile liner strain 1000 1000 1000 9
" 3000 j (micro infin) _ _ _ _ _ _ _ _ _ .- -- __
Maximum compressive liner strain 11001 15601 5701 22671 III (micro in/in) i' Allowable compressive liner strain 2000 2000 2000 5000
, (micro in/in) ~
i, , I Pressure at which first concrete cracking l No Thru No Thru No Thru 631 occurs through section thickness (psi) Cracking Cracking 2 Cracking 2 l IV
! Maximum pressure for transient (psi) 0 0 63 83 '
U
, E I Tensile strength of concrete in cap taken as zero. Cap reinforcement has practically no effect on h dome stiffneas. Therefore, values are practically same as for 24" dome. -
y 2 Due t o assumed zero concrete tensile strength, cap cracks at 0+ psi pressure. Remaining 24" dome i i U thickness is sufficiently prestressed to preclude cracking.
O m
s Referenced to unstressed condition. 5 Due to 1.5 Pa+T a "Due to 1.15 Pa i
d i Appendix G Cont'd. Page 8,of 9 4
-3 .
1
SUMMARY
OF RESULTS
.1 STRUCTURE: 36-Inch Dome, Italuding New Reinforced Cap i
4
! Dead Load Norm. Winter LOCA 4 Operating Structural @ 40 Years Item Prestress Condition Integrity No. Description @ Early @ Early Test Plant Life Plant Life
},
Maximum concrete membrane compressive 3 l
, stress (psi) {
Maximum concrete membrane tensile -5522 4453 stress (psi) j Allowable membrane stress (psi)
- (compression -/ tensile +) -2700/0 -2700/0 -2700/0 -2700/0 t - - -
i'
<_,w s- _.. _m. . _ . - - _ _ _ _ - - _ . _ . . - _ _ - _- - -- - -- - - - -
i Maximum compressive concrete extreme fiber stress (psi) - Uncracked analysis value 3219 3999 +
a 1- i' i Cracked section investigation value -
VI Calculate when actual uncracked is Not Reg'd 354 0'+
greater than allevable '
Allowable compressive extreme fiber g,
3600 3600 3600 3600 ,
-- - - - - - - - - - -= ~==,~----w-~
-n- __
Maximum concrete tensile extreme fiber T stress in areas without reinforcement k
e (psi)
VII Allowable concrete tensile extreme fiber 0 0 0 0 stress in areas without reinforcement T (psi)
M 4
12712 occurs in bottom 24"; corresponding cap etress is 3 Cap reinforcement adde.d to resist all net membrane 365 psi tensile. (Sta 161.1 9). tension. See Figures 5-20 and 5-21.
2No tension exists in bottom 24" section of dome; minimum compression Indicated. Cap is cracked; reinforcement at 11000 psi tension, Occurs at Sta 150.2 I i
i y
j Appendix C Cont'd. Page 9 of 9 i
SUID!ARY OF RESULTS STRUCTURE: 36-Inch Dome, Including New Reinforced Cap
$ Dead Load Norm. Winter LOCA
+ Operating Structural 0 40 Years Item Prestress Condition Integrity .
No. Description a @ a Test p
i Calculated shear at point of minimum 45 55 6 10 margin (kips /ft)
Allowabic shear at point of minimum VIII, margin (kips /ft) 55 55 7 23 4
Location of minimum shear reinforcement 151.5 152.0 161.0 160.0 margin .
_ _ . ___.. . . _ _ _ - . - _ - ~ ~ - - - - - - - - - - - - - -- - - - - - - - - - - -
- = - - - -
p 54000" l
- Maximum , tensile stress in reinforcing 748 2 ,3 70312,3 110004 @ Sta 161.1 bar (psi) (Fig. 2)
IX . ;
Allowa.,le stress in reinforcing bar 200003 200003 30000 4 54000 4 (psi) ,
1Since cap is cracked af ter SIT, shear results are same as those for 24-Inch Dome.
! 2 Practically the same as for 24-Inch Dome, g 3In location where Grade 40 reinforcement exists.
N 4In location where Grade 60 reinforcement exists, 0.5fy for SIT and 0.9fy for IDCA.
{D e
f I to b
t
+
. .. ... t APPENDIX H CADWELDING REQUIREMENTS Revised: 9-22-76 s, y,. .y,,7 . ,
. . - . ~ . -- -_- -- ._ _ _ . _ . . __
. 4 APPENDIX H CADWELDING REQUIREMENTS 1.0 Applicable Codes, Standards, and Reference Documents The WORK shall be in accordance with applicable portions of the following codes, standards, and reference documents:
- 1. American National Standards Institute (ANSI) N45.2.2-1972, "Packagir.g. Shipping, Receiving, Storage, and Handling of Items for Nuclear Power Plants."
- 2. American Society for Testing and Materials (ASTM):
- a. A 370-75, " Standard Methods and Definitions for Mechanical Testing of Steel Products."
- b. A 513-73, " Electric-Resistance-Welded Carbon and Alloy Steel Mechanical Tubing."
- c. A 519-73, " Seamless Carbon and Alloy Steel Mechanical Tubing."
~
- d. A 615-68, " Standard Specification for Deformed and Plain Billet-Steel Bars for Concrete Reinforcement.
2.0 Cadweld Splices
2.1 General
Splice samples shall be sister splices (removable splices made in place next to production splices and under the same conditions).
, 2.2 Cadweld Operator or Crew Qualification:
- 1. Prior to performing the actual splices of reinforcing bars each member of the splicing crew (or each crew, if the members work as a unit) shall prepare two qualification splices for each of the bar sizes (identical to those to be used in the structure) to be used in the production work. The completed qualification splices will be visually inspected and ter.;ted for tensile strength.
- 2. Each member of the splicing crew (or each crew, if the members work as a unit) is subject to requalification if a) splicing has not been done for 3 months or more or b) completed splices fail to pass the visual inspection test or fail to pass the tensile test requirements or c) there is any reason to question their ability. The requalification procedure shall be identical to the original qualification procedure.
H-1 Revised: 9-22-76
,,,,, m m m ..: -- -- -- - - - - - -- -- -
l'
(~
APPENDIX H (Cont'd) 2.3 Acceptance Criteria for Cadweld Splices:
- 1. ' Sound, nonporous filler meterial shall be visible at both ends of the splice sleeve and at the top hole in the center of the sleeve.- Filler material is usually recessed 1/4 inch from the end of the sleeve due to the packing material and is not considered a poor fill.
- 2. Splices which contain slag or porous metal in the riser, tap hole, or at the ends of the sleeve shall be rejected. A single shrinkage bubble present below the riser is not detrimental and should be distinguished from general porosity as described above.
- 3. There shall be evidence of filler material between the sleeve and the reinforcing bar for the full 360 degrees; however, the splice sleeves need not be exactly concentric or axially aligned with the reinforcing bars.
1
- 4. In order to qualify, the completed splices shall also meet the acceptance requirements of Erico Products " Inspection of the Cadwald Rebar Splice," Standard RB5M-274.
- 5. Splice samples shall be subjected to tensile tests by using loading, rates _ set forth in ASTM A 370 to determine conformance with the following acceptance criteria:
- a. Individual splice strength criteria:
t The tensile strength of each sample tested shall be equal to or exceed 1257. of the minimum yield strength specified in ASTM A 615 for the grade of reinforcing bar being used.
- b. Group splice strength criteria:
The average tensile strength of each group of 15 consecutive samples shall equal or exceed the guaranteed ultimate strength for that grade of reinforcing bar as speciffed in ASTM A 615.
2.4 Positioning of Reinforcing Bars for Sister Splices:
Sufficient ~ extra lengths of reinforcing bars shall be positioned in those areas of reinforcement where splices will be required prior to any splicing taking place in that concrete pour. The reinforcing bar shall be secured in position so that it cannot be displaced during splicing. The positioning of the reinforcing bar shall be such that the location and orientation of the sister splice are similar to the splices it will represent.
H-2 Revised: 9-22-76 m,wgame.wrw.a ? ~~ - ~
v- - -- - - - -- - - - - -
P' APPENDIX H (Cont'd) 2.5 Availability of Sample Test Results:
Test results for the samples shall be available and be certified as being in conformance with the requirements before concrete is placed.
3.0 Frequency of Testing 3.1 Separate test cycles will be established for splices for each bar size and for each splicing crew.in accordance with the following frequency:
- 1. One sister splice of the first ten production splices.
- 2. Four sister splices for the next 90 production splices.
- 3. Three sister splices for the next and subsequent units of 100 splices.
i H-3 Revised: 9-22-76
~ m ,p u v m -e en ~ = , -
.- - - , . . . . - ~ ~ - - o-- - - - - - - - - - -
N APPENDIX I RADIAL TENSION AND SHEAR Revised: 9-22-76 y s,c y-s,, , . . ,- -= - - - -- - - - - - - - - ~ --
s APPENDIX I RADIAL TENSION AND SHEAR GENERAL Fundamentally, radial tension is directly related to compression stresses in I the plane of the dome. The importance of considering these stresses in l conjunction with radial shear stresses is dependent on the load condition being considered; i.e., service or abnormal. l For abnormal load combinations, where in-plane tension exists at the top )
surface, radial tension stresses become compressive. Therefore, radial shear stresses are considered separately.
For service load combinations where cracking does nor occur, radial tension and shear stresses should be combined. Mohr's Circle graphical method is an appropriate technique for accomplishing the combination.
The critical location in the existing structure for radial tension and shear in combination is in the pours outside the main delamination; e.g. , pours G and H.
Those pours are reinforced (original design) through the thickness of the dome with #8 (]) reinforcement (see Figure 2-2). The minimum area of steel provided
,e perpendicular to the plane of the membrane is approximately 0.79 in.2 of steel per 432 in.2 of concrete.
SERVICE LOAD To illustrate the combination of radial tension and shear for service loads, an example state of stress for the original structure will be evaluated. Considering the middle of the outer tendon conduit as the critical section for radial stress, an average value of 40.8 psi of radial tension stress is definsd in Figure 3-15. The state of membrane compression stress compatible with that value is approxinetely 1800 psi (original design, see Figure 2-12). To define a
. complete state of stress, a radial shear of 50 kips /f t (from Figure 2-14 at analytical station 1540) or 116 psi nominal shear stress for 36" thickness is used for example purposes. This shear is larger than the shear that occurs at the same location in the structure as the radial tension and compression stresses. Figure I-1 indicates the principle tension stress is 48.1 psi and that plane on which it occurs is approximately the plane of the membrane.
Therefore, the steel stress (for each #8 bar) is, f, = 48.1 psi x 432 in.2 26,300 psi 0.79 in, which is 0.48fy for f y= 55,000 psi (mintnum mill test result) and is acceptable.
This steel stress is an upper bound solution for pours G and H (original, delaminated and repcized structures) since the combined membrane compression,
, radial tension and radial shear state of stress anywhere in those pours is less
\_ critical than the example case.
I-1 Revised: 9-22-76
.-,-.c _ - . - _- .- -
[' FACTORED LOADS Radial shear capacity (as a measure of diagonal tension) is of importance in factored load considerations to preclude premature failure. Each of the governing load combinations has been investigated for the existing structure
. with regard to shear capacity required for factored loads (see Figures 4-6, 4-12, 4-18 and 4-27).
k I-2 Revised:9-22-76
.n ,---e sc. . - --- -- - - -----
n 4 o.S ISoo J 1800
" IIG y 40~.S 40.B PSI 5%TE OF STRES5 tn 1 ,
W !
t9 . c.
= 1 o [ $ !
l' '
~
y J ,
,. n 1
' \ f t a
\
\
\
4 8 .1 P S )
1800 PSI -
MOHR'S CIRCLE Cousisao STaesses F I G U RE. 1-I Revised: 9-22-76
- g. v s Wqn -
sy< pp W ""
. ~~
v^,
Y '*W*M ""W ^"W***"'""*' ~ * " ^'
r
('-
APPENDIX J NRC CORRESPONDENCE
/
SIT INSTRUMENTATION
(
Revised: 9-22-76 7'
J f. y iF ,
y -" .h '
- J -
APPENDIX J NRC CORRESPONDENCE SIT INSTRUMENTATION September 1, 1976 TO: MR. LEON ENGLE FROM: F. L. MOREADITH THE ATTACHED LETTER IS SENT TO YOU PER INSTRUCTIONS BY MR. B. L. GRIFFIN AND MR. J. T. RODGERS AND DESCRIBES THE EXTENT TO WHICH INSTRUMENTATION DESCRIBED IN THE FSAR WILL f- BE SUPPLEMENTED AS A RESULT OF THE CRYSTAL RIV$R #3 DOME REPAIR.
SHOULD YOU HAVE ANY QUESTIONS CONCERNING THE DESCRIPTION, PLEASE FEEL FREE TO CONTACT ME.
FLM:cd 4
s J-l Revised: 9-22-76
' n ,~-.wmm n ,; - -n, . , , nn. -.-m-: , - - , - - - - .-- - -
~
_ .. ..._, . . . _ ...m ._ . _ _
l l
l i
Gahert/commemmasseh w. , _ 1
.(
easse assoceu sc. e a a uma n a a see I
APPENDIX J (Cont'd)
August 31, 1976 FPC/DR-52 Mr. B. L. Griffin Vice President and Manager of Crystal River Unit #3 Dome Repair Florida Power Corporation P. O. Drawer 1057 Crystal-River, Florida 32629 Re: Crystal River Unit #3 Dome Repair Instrumentation
Dear Mr. Griffin:
\
The Containment Building's structural response to the SIT will be monitored '
- - by utilizing the existing instrumentation described in the FSAR and supplementary instrumentation installed within the repaired dome.
Instrumentation within the repaired dome will consist of a series of instrumentation stations established in the dome on two (2) orthogonal axes at distances of 15, 30 and 45 feet from the apex (see Figures 1 and 2).
The data recovered'at these stations consist of the following:
~
- 1. The hoop and meridional steel strain changes on the liner using SR4 three-element electrical resistance strain gages attached to the inside surface of the liner.
- 2. The hoop and meridional concrete strain changes near the upper surface of the lower portion of the dome using Ailtech Concrete Embedment Gages with a 4-inch gage length.
- 3. The hoop and meridional reinforcing bar strain changes of both
-layers of steel within the new concrete cap. These measurements
'are obtained from SR4 linear strain gages attached to #4, Grade 60 sister bars.
.(
, J-2 Revised: 9-22-76
~
- m _ ~ e n .- m ; ,,
_w. . - - - - - . - .- -
%C' APPENDIX J (Cont'd)
Mr. B. L. Griffin FPC/DR-52 August 31, 1976 Page Two
- 4. The reinforcing bar strain changes of the #6, Grade 60 radial anchors. The gages are SR4 linear strain gages.
Gross structural deformations will be measured by extensometers attached to the steel liner plate at the locations identified on Figure 3. This instrumentation provides the following data:
- 1. The vertical displacement of the liner, inside surface, measured at three (3) azimuths and at radial distances of 0, 29, 49 and 56 feet from the apex.
- 2. The radial displacement of the liner measured at the 49 and 56 foot radii.
All instrumentation readings are processed and recorded using a Vidar
( Model Autodata 8 data acquisition system. The data sampling rate will be sufficient to adequately evaluate the structural response of the structure. '
Sincerely yours, F. L. Moreadith Project Manager Crystal River Unit #3 Dome Repair TDB:cd cc: B. L. Griffin J. Alberdi R. S. Burns C. E. Jackson J. E. Colby A. L. Gomez i
I
(
l J-3 Revised: 9-22-76 l
, wme m,, n , n-, - .- -
,n _. n . - -
App:ndix J Cont id. -
O*
-- 15' RADI U S
(:--
TO INSTRUMENTATION GROUPS I
45:o.,
?IS.'g' 30'.o a
_9 O* '2.70
\
i l
l l
I B O*
Pt_A N STRA\N INSTRU MENTNilON FOR' MONITORI NG STRU C.7URM-. RE S PON S E DUR\NG SIT I J-4 Revised: 9-22-76 l
~
Fic,iiru '.
p w q w gr v--*gw- w
Appindix J Cont'd-I
- ll d
MERIDION AL i HOOP STEEL #BMERIDlONAL (Hoop STEEL
\ Nl s.- / hw g j I . _p . /Y / y U
G U^ f( : *.' -
':U 3 o'
. : . . '.- g o l
g on n 1 4.,.- Z 2 0
f \f l i s! :*?.~f V f
l .;}.;"-
L-t#4 SISTEF! BARS DICO SERIES h WITH SR4 LINEAR TENDON W '
STRAIN G AGES DUCTS [
~ % RADIAL ANCHOR k'2.- DIRECT 1O N AL E WITH SR4 LINEAR AILTECH CONCRETE 9 2 r
STRAIN G AGES EMBEDMENT GAGE y 0 I
._ STRAIN GROUT U.'
.\ -
.. . 4: ,
k" LINER PL -
. k .' [i;,'-[
( s'-i ' .i _ : 4 -: V I
3 ELEMENT SRA. ELECTRICAL RES\ STANCE STRAIN GAGE TYPlcAL SECTION AT IN STROMENT 4.T IO N LocATi oN s
1 Fi c,une 2. l J-5 Revised: 9-22-76 ;
-, -r
.-. . ~ - . . . . . . . . _ . .
App ndix J Crnt'd.
APEX
/ \
l 4"SPR AY RING i
LEL. ?.70'- 8"
{EL. 267' G" _ ?.8'- 8h"_
_ 4 9'- 3 ii." a
- 5 5'-8 $"
i LOCArlON DIR. 0F ME ASMT.
AZ 90* EL. ?_7 0'- 8" ' RADI AL 1 A2 200* l l
~
l AZ 333*55' i AZ 90* E L. 2 67'- G" AZ 200* l l AZ 333*55' t i AZ 90* 4 SPRAY R\NCi VERTlCAL l E L. 2.70'- 8" l t E L. 2 GT- G" AI 200* 4N SPRAY RING L i l E L . 7.7 D' - S" t E L . 2 G7 '- (o" A Z. 333*55'47" SPRA,Y RI NG l E L . 27 O'- B" l f
~
E L 2 G7 '- G" l DOME APEX t
.E xTEa soutTsas Os Liwa w Foa Dispusc.emasT Mt Asu Ramaw T J-6 F1co at 3 Revised: 9-22-76
t MP sguest aus suuuages
- - [ . .'
assem assuevis sc. e o e me anse um
,4PPENDIX J (Cont'd)
SEPTEMBER 14, 1976 FPC/DR-75 MR. LEON ENGLE-DIVISION OF PROJECT MANAGEMENT NUCLEAR REGULATORY COMMISSION WASHINGTON, D.C. 20555 RBi CRYSTAL RIVER UNIT #3 DOME REPAIR INSTRUMENTATION RESPONSE TO C. P. TAN'S TELEPHONE CALL'0F SEPTEMBER 13, 1976 DEAR MR. ENGLE:
IN RESPONSE TO DR. TAN'S TELEPHONE CALL, THE FOLLOWING ADDITIONAL INFORMATION id PROVIDED WITH REGARD TO INSTRUMENTATION DESCRIPTION.
THE EXTENSOMETERS BEING USED AS PART OF THE INSTRUMENTATION FOR THE CRYSTAL (k RIVER UNIT #3 DOME REPAIR-AND SIT CONSIST OF A TRANSDUCER ELEMENT WHICH IS AN INFINITE RESOLUTION LINEAR POTENTIOMETER. OUTPUT OF THE POTENTIOMETER IS A VOLTAGE THAT IS PROPORTIONAL TO THE DISPLACEMENT OF THE MOVABLE CONTACT (SEE ATTACHED FIGURE) WITH RESPECT TO THE POTENTIOMETER. MOVEMENT OF THE 4
LINEAR POTENTIOMETER IS ACTIVATED BY INVAR WIRES ATTACHED TO THE LINER PLATE AND THE MOVABLE CONTACT. THE INVAR WIRES ARE REPT UNDER CONSTANT TENSION BY CONSTANT LOAD LINEAR SPRINGS.
SPECIFIC CHARACTERISTICS OF TEE EXTENSOMETERS BL dG 1!MPLOYED ARE:
MOVEMENT RANGE 1.75. INCH RESOLUTION 0.001 INCH l
ACCURACY 'O.005 INCH LINEAR SPRING FORCE 20 POUNDS 1 TO ASSURE THAT THE MEASURED DEFLECTIONS ARE DUE TO MOVEMENT OF THE CONTAINMEFT STRUCTURE AND NOT THE SUPPORTING STRUCTURE THE FRAME OF THE TRANSDUCER UNIT IS ATTACHE 1s TO A " FIXED" REFERENCE STRUCTURE. ATTACHMENT STRUCTURES FOR THE TRANSDUCER UNITS WITHIN THE CONTAINMENT BUILDING ARE THE TOP OF THE SECONDARY SHIELD WALL AT ELEVATION ~180'-6" Ab' THE OPERATING FLOOR AT ELEVATION 160'-0".
THE EXTENSOMETER TRANSDUCER FRAMES HAVE BEEN.DIRECTLY ATTACHED TO THESE REFERENCE STRUCTURES OR TO SPECIALLY PROVIDED STRUCTURAL FRAMES ATTACHED TO ,
- THE REFERENCE STRUCTURE. l N i t
J-7 Revised: 9-22-76' I
- i. .
l 1 $ A M-r-- QT , .
- i' ' " *
.~
.3 .. ,
___ _ _ "_" ' ' _. "_"7 f
.(
APPENDIX J (Cont'd)
MR. LEON ENGLE NUCLEAR REGULATORY COMMISSION SEPTEMBER 14, 1976 PAGE TWO IF THE ABOVE INFORMATION IS NOT FULLY RESPONSIVE TO DR. TAN'S REQUEST, PLEASE CONTACT US FOR FURTHER INFORMATION.
SINCERELY,
?Y F. L. MOREADITH PROJECT MANAGER CRYSTAL RIVER UNIT #3 DOME REPAIR
- FLM
- cd ATTACHMENT cc: B. L. Griffin
, J. T. Rodgers W. R. Zimmerman y
A J-8 Revised: 9-22-76 s.mm,-. . - , - - , _ ~ . - . _ _ . . _ . _ . __ .
__m.__._,. _ _
'sv - - _ - e , --r.. . - - -o---
__.~.. _ -.. - --
Appendix J Cont'd.
{ s . ;,: l
. :;g U NATTACHMENT PL WITH EYE BOLT WELDED To LINER 1
f -INVAR WIRE E% TEN 50 METER
'SEE DETAl L *A" TYPICAL DETAIL Oi=
Ey.TENSOMETER INSTAL:.LATION
= -lNVAR WIRE N
TURNBUCKLE MOVABLE CONTACT i-i 1- -l gCHANNEL FRAME k ,
I LINEAR - " I81 SPRINGm"'-M'y Mll '
INFINITE RESOLUTION l LINEAR POTENTIOMETER LEADS TO DATA Mj '
AC.QUISITION SYSTEM % /
1 gjv; [ ATTACHMENT FLOOR OR OTRER To
,pg,AEO STRUCTURE f
DETA,i t " A" J-9 FIGURE l Revised: 9-22-76
' ~ ' ' -
. . _ 'i _-_7,_
--z- .--, _ J- -
9/15/76 3:40 pm APPENDIX J (Cont'd)
( -
Mr. Engle called wi th the following:
Mr . .T . C. Rodgers:
Re: SIT Instrumentation Crystal River #3 Dome.
h*e have reviewed the in information on SIT instrumentation for the dome as contained in the September 1, 1976 telecon from F. L. Moreadith to Leon Engle.
It isourunderstandingthe*"ff"* "which are for measuring the deflections at pertinent locations of the dome will be used in combination with strain gauges an concrete, reinforcing bars and steel liners.
h'e concur that the proposed instrun. station if p;operly installed should provide information required ror adequate assessment of this structura.1 int.egrity of the $ repaired dome.
Unless you have some questions pertaining to the telecon, you should consider "go ahead" on SIT instrumentation for CR#3.
Leon Engle
/Y' M AM"
/ d<
/ ,,
d& . -
s
.p_ zz sxe c-aZ - - --
3-1o
= -
. 1..e. ..m., e t
.3,, r ,
EQ H o n.6. h6? O &
n
(* uv~m$n
'*?h~e -
()VYIDI C0p*0ma* Og August 10, 1976 United States Nuclear Regulatory Commission Washington, D.C. 20555 Attention: Mr. John F. Stolz, Chief Light Water Reactor Branch No. 1 Division of Project Management Re: Florida Power Corporation Crystal River Unit #3 Docket No. 50-302
.c Gentlemen:
In accordance with your letter of July 30, 1976 relative to the request for information concerning our interim report, " Reactor Building Dome Delamination," June 11, 1976, we submit the attached e response as supplement number 1. Forty (40) copies are included.
We have prepared this respor..a for submittal at this time per your request as a supplement to the interim report and in the next two weeks we will lorwar,i the agreed upon correction / addition pages to amend the re* port to taflect the repair procedures being followed subsequent to July 27, 1976.
It is anticipated that the format being followed will allow a single report to suffice in furnishing the necessary information for you and your staff. Additional supplemental pages will be submitted as required by your review.
We will discuss this matter with Mr. Engle today after our presentation on the repair status. O Very truly yours, 1
b J. . Rodgers Assistant Vice President.
JTR:cd cc: Mr. Norman Moseley
, Atlanta, Georgia A
Region II Inspection a Enforcement vp ok'N4 4 ^s i
~-
APPENDIX K DISCUSSION OF CONCRETE AND AGGREGATE PROPERTIES 1
I l
l l
s-l l
l Revised: 12-10-76 memsw.v wpy-wc-m- rc ~ ~ - - + - ~ ~ - - - -- - ------ -
APPENDIX K DISCUSSION OF CONCRETE AND AGGREGATE PROPERTIES
.: 3, ;
,. .,. l 217 356 4398 R.F.D. 3, Champoign, Illinois 61820 i;ovember 24, 1976 Dr. Frederick L.11oreadith Project lianager, Crystal River Unit.3 Dome Repair Gilbert Associates. Inc.
. P. O. Box 1498 Reading, PA. 19603
Dear Dr. floreadith:
RE: Crystal River Unit 3 Dome Delamination The cause of the delamination that occurred in the dome of Crystal River Unit 3 presumably during prestressing, is not clear but is most likely related in a major sense to the quality of the aggregate. ,
The coarse aggregate was obviously weak although it met what were considered, by the owner, to be the pertinent ASTM C33 requirements. The low strength of the coarse aggregate is indicated by:
- 1. a loss of 41.4 percent in the Los Angles abrasion test, even though this was within specification limits,
- 2. the nature of the particles, many of which were fossiliferou.:
with high porosity and permeability, and
- 1. the fracture of all of the coarse aggregate particles at the delaminated surfaces Normally, concrete is a more ductile material than cement paste because the aggregate particles act as crack arresters and more energy is required to propagate a crack in concrete. For the concrete in question, the fracture resistance of the aggregate was probably less than that of the cement paste; certainly it was no greater. Concrete of this type would be expected to exhibit, comparatively, a more brittle and sudden failure than concretes made with good quality aggregates. Also, the tensile strength would be expected to be low and was measured in tests at frci 230 to 505 psi.
The average tensile stress in the dome concrete as determined by Gilbert Associates is 41 psi, or perhaps twice this amount if allowance is made for the area occupied by the tendons. However, a finita analysis indicated high local tensile stresses in the concrete at the interface between the con-crete and the tendon ducts by, I believe, assuming the concrete to be a 4
. linear elastic material. Had the dome been made with concrete containing good quality aggregates it would not be a linear elastic material and the j
. maximum tensile stresses would be considerably-less than those computed.
v K-1 Revised: 12-10-76 1
nw s,.- wan - - - --- - - - --- -
- l
. . - . ~ - . . _ - .. . -.
APPENDIX K (Cont'd) f Dr. Frederick L. Moreadith fbvember 24, 1976 Page 2 The maximum-stress would also be influenced by thrinkage, creep and microcracking. Furthermore, concretes by their very nature are abundant in flaws, cracks and irregularities which are all stress raisers.
Therefore, the effect of the conduit in a dome made of concrete with good quality aggregates may be nil, certainly, it would not be the effective stress raiser that the elastic analysis approach would indicate.
However, the concrete in the dome of Crystal River Unit 3 was made with a weak coarse aggregate. These aggregate particles apparently did not act as crack arresters in the concrete and thus tre concrete was probaoly as brittle, or 1early so, as cement paste. For this reason the elastic analysis approact suggests a source of high stress which could initiate cracks. This is not to say that the stresses are as high as computed but a fraccion of their computed value is sufficient to cause failure in this particu'ar concrete.
If there is. merit to this approach and the dome concrete is as brittle as speculation indicates, little energy would be adsorbed by the dome concrete abaad of a crack and a crack could propagate with relative ease.
The weak coarse aggregate particles appear to have had a compound effect: (1) they increased the brittleness of the concrete which in turn resulted in an increased maximum tensile stress, (2) they resulted in a concrete with low tensile strength and (3) they permitted cracks to propa-gate with low energy input.
While more data is needed, it seems that although delamination of the doine occurred when this particular aggregate was used in the con-
, crete it would not have occurred had good quality aggregate been used.
Sincerely yours, Clyde E. Kesler l CEK:cds l
N- ,
l K-2 Revised: 12-10-76 {
- , , , _ .. . . - . , _ m - - - , - - -
i
+r SUPPLEMENT 1 RESPONSES TO NRC COMMENTS TO THE CRYSTAL KIVER #3 REACTOR BUILDING DOME DELOfINATION REPORT DATED JUNE 11, 1976 DOCKET NO. 50-302 GENERAL
- 1. For. easy reference, provide a list of tables and figures in the Table of Contents.
Answer: The material suggested has been incorporated into :he Table of Contents. [
s ,
S-1 Supplement 1
- Revised
- 12-10-76 i
k [TidWWr , gw- . - - - - 3 %,$~ D"g'; I4QN*
- W T* * * * * ,y"% www+= ^ - p-.. . , . m,,,...
. .. - . . . . .-=.
SECTION 1.2
- 1. The staff considers the establishment of th= causes of the dome delamination to be important in assessing the adequacy of the repair program and in providing assurance that another crack will not occur again during the life of the structure. The potential contributing factors should, therefore, be identified indicating the magnitude of radial tensile stresses created in the concrete.
Answer: The material has been incorporated into the report in Section 1.2.
- 2. The use of radial anchors will enhance the capabilit'y of the dome to resist radial tension. However, they will not eliminate tension in concrete, and therefore small cracks may still exist. Provide an analysis to indicate that such cracks will not jeopardize the required structural integrity of the dome to resist all combinations of loadings for which it is designed.
Answer: These cracks exist primarily in regions where membrane behavior dominates, i.e., negligible shear stress across the cracks.
Despite the presence of the cracks, the membrane compression capacity of the concrete is adequate. Under LOCA or SIT, there is radial compression across the cracks.
It is of interest to note that under 15% detensioning (admittedly a nominal stress change) deflections were less than predicted. If the cracks were contributing any significant effect to the response, larger rather than smaller deflections would be expected.
In addition to the above considerations, secondary cracks will be epoxy grouted and the radial reinforcing will cross virtually all these cracks.
l
( S-2 Supplement 1 8-10-76 M %@% Th% %M$MRN **u '
.=5fMMM%F Merr #[ '7 F u. -' ~ ~ D"# ~ ~
~ ' * '
l l
f~
SECTION 2.3 AND TABLE 2-2
- 1. Clarify the definition of tensile capacity of concreta. Explain how principal tension is related to shear d diagonal tension as indicated in Section 2.3.1, and what is the differu ce between the shear discussed in this section and that in the next section (2.3.2).
Answer: The material in sections 2.3.1 and 2.3.2 have been rearranged in the report, section 2.3.1 title is "Tlexural and Membrane Tensile Stresses" and section 2.3.2 is entitled " Shear".
- 2. Provide and describe with examples of actual design, the conditions under which each of the criteria (a) and (b) in Section 2.3.1 is applied.
Answer: The material requested now appears in section 2.3.2. Attachment #1 is a design example.
- 3. Since the stress / strain distribution is tri-axial, the limits of 3 /f -andc ^
6 /(( may not be directly applicable to this problem and their use should be justified.
Answer: The state of stress in the dome may be regarded as being biaxial since
, the stress in the radial direction is very small in comparison with the 4 \
membrane stresses. The interactions for tension-tension and tension-compression are not significant at least until the compression exceeds about 60% of the compressive strength of the concrete (Kupfer, Hilsdorf and Rusch, ACI Journal Aug.1969) (Ref. 8 of the Report). Thus the limits of 3 /(( and 6 /(( are justified.
- 4. If 0.85ff as extreme compression in ultimate strength design is used, it may not be directly applicable for the same reason as in the above comment and should be justified.
Answer: Although criteria indicate that under factored load concrete stresses would be allowed to reach 0.85ff they do not. The actual stresses are much lower and do not appear critical since the dominant stress is bi-axial compression the strength should be higher.
- 5. The shear strength of concrete is influenced by stresses orthogonal to the axis of the element; therefore, this effect should be considered.
Answer: Hoop tension stresses should have little or no effect on radial shear strength, since sufficient bonded hoop rebar has been provided to preclude hoop tensile " Failure".
S-3 Supplement 1 8-10-76
- JA (1, Erf,v w -7y - . - pea - ( %* , b=, d [I@ D k.[1, [;
C 5 ..
'~
~
Shut 1 of 4 ATTAC10fENT 1 FOR ANSWER TO QUESTION 2.3(2)
/V2C $tvs?Ior) 2.2.2- Shear Lhsian Emmoles dace- e. Membrane /ension orincinbane comp. < /copsi Emmple LocA 26" t%me (Fig 2-30 f 2-se) 6fe.Is2.s*
Fmm .she// analysis V = 9 4 "'/s Nip /72 Y(7e"n of) p = 98fai (comp) n::inbare
~Since 6=q =98pa;(cmp) <: toops; Leomp),.uce sie-g s, ch 17.
Calculata conc. capacity usny (17-2) v.a4 ( n-3).
4 A r d 4 ,. a, ainee no =-);ese rein. avail. a f is,.s .
s .+
Froin (17-s), .
M'=M-W(4$ci =
= /22.6 f 6 172-42.3 (4936-318 gypg *
/v'= 98p; v 12u ss". 4 2.3 */i 2 = ss" ., -
31 8* ( di d . fa m r p to of iek) ;
1 froir, (n-2)> l Ge + (I.9Tl6' +2sooj.% \/c/ ) (,0w h = i;*, g,.g
- m, 00*) i
\
'l/c = o.8G h 9Tsteo + 2sco (. cob)(94)(si.eij'/22.s]
% - 129.8pei (< s.E+ %fl e ziopa) l
\
\lc = bd tL = (12)(3/.8)(129.8) = 49.5 "A 5 2-32
/CCD 9 L
U/Aw> ate. appliec/s/xar, Vu., -= V= 9 4 */i l 49.5 "/i c:cp2ci4j '.> 9 4 "/i .~.' s s A k o.e IC 2 -3 =
9 Supplement 1 8-10-76 1
[eh y kg s'.8h O ht*-- '$. PgJ-5, e e g g ") YY * #
4N O% %# g -w
Sh ut 2 of 4 l
ATTACHMENT 1 FOR ANSWER TO QUESTION 2.3(2) (Cont'd) Supplement 1 -
A/EC Quedbn 2. 3- 2 Shear Des 13n Exainple*" s Case b. - Membiane Gmo<ssen > !%i .. . l 2y3mp/e hl0!**/Y1. YllAb*r 0 Cr. Conc l. 6" Od!Hd' 6 $~l0f2"E#h sta. . t es. s ' :
Pr~ein she// analysis.
V = 1/9 3 "/i N)p = 277 l(Ten o/-) (74= /384p .i (coinp)
/Yleiribrese chess /284 psi camp y toopsi cornp
. . use sis - G B ~ Ch z 4, ( P/s f.onc)
My $ .dbow ok t/cii mininxim concreta. capaci6 , d Va , p .us l shear sein. ' 6Av /b-l e _,,
If v L,, -i c6 A-s4 2 %. > oe 3 l s
ifnof calc \/e f,,,n (es-iz) vvi/h FsM mcep%s.
l t
+ Vemin 4 (17Ce/)Nd' ,
' . el = /8 4"(elisi fam ed,ene
=
c.mp. Asec & eg +:netons)
<;6 Wmon22,56o =(o.85)*/t0 7)G2Kte.4Necoo e 2 2 G */s cor)cte/c.
- Av cl4 . (o.se](0.,<1)(si.8)(40ksi) .189 8 *// sh-ar iein.
S <;t. S .
l d - def. (l,,.n e,4,eme comp Ak b /ension ein.
s = 4Yz' meric/iona/ spacing o /* # 2 :sk ar bars (i2" hoop .: spa)
. ' - Conc + fein : '2l'2 4 "A Cepac.ify Applied ({achad) > % = l SV= /.5(II9 3 */-) = /Icf.o"A
- '. < : -: '2l2.4 */i &a28cihj > /79.0 "/r app /ted(f;,ek./), .'. .
S ec fiori is OK basec/ en minunuin conc shsai- capacity
Shnt 3 of 4 -
ATTACHMENT 1 FOR ANSWER TO QUESTION 2.3(2) (Cont'd) Supplement 1 8-10-76
,. /VEC Gye.41an zE-2 :Thesr Desen 5en:ple:
6ese b.-/ Adenshi=ne Cemw.uisn > imi>=.i E a m p le A/ erin. win 4r oper. Cond. 26" Leme (-Tp c-is f2-20) da /52. o
- F<cin ,=rhe//analycis
% 46.Sk// tvl9 207 Y(Ten of). v% = l48Epr.1 (com,o)
\/q =/.S\/= /.5(45.5) = 68. $ 4;/s Memb.cne .:stww = idespaicemp > ion;ci< i mp
.'. us e El8-4.5, Ch zG (P/s cct;c) sh trJir) \lc f .:Chsa.r. tein (# 8618v7H)
& \lc min '2'2.G </t (oIdsnied p1Ntous/3 )
16 Av dk , (0.85)(.79/t.s)(21.8)(4s)
~
3). 8 */
- s '/ 8" Ma.f c a p. = 2 2. 6 + 5/. 8 = E4 4 '. 'r < S 8 8 4/i cale actus2 Ve ;%n (es-iz) w,tese 4=pnoru
\lci = Esv b'elN f' +. ,C O 6 + V s'
~
.sl z
Es.v = I.75'- 0 M h
+- 4 of 7
Mcr = I (G W + fpe ?$)
U
{pe -/g~= comp,essive shess in concan due b pus-hecs end deid load cZth :all l'=eseo, a/ //>e 9fieme Abu af' ele arim.
a t whkL -lenc. ion ,cheases a>< causad
% cpplied loads .
M & 'v' eut dita. $ a// /oac/s d = te.4"(disf. from ec/ieme comp. A ber- 70' c9 7E'ndons.)
i
ATTACHMENT 1 FOR ANSWER TO QUESTION 2.3(2) (Cont'd)
Shast 4 of 4 ,
$8CC b.~/ (CanW)
'~
, dlnce h'lf c h t.: 4 /$ nsiori o 77 o5, ,0=ff , *
,0 =fof = *fhd=I*'f/2)(3/.8)c 0.0042 f) = 24/4 = 7 25 Esv = !.75 - 0. 03G _ + 4(7 2cVoca2) = 0 69e > 0 40 (725)(.0042) Uc.d. O.60 I= k QC)(36)'= 464GYoi,,4
> ef = Es/1 s /?"
d Vf*' a G'Ys0*o a 424p.si Spe -(J = --9/o pai camp (D + f : -9to of, - 2soo zf)
.e Nict = 46se6 (STsooo +<1to)y
/ gi /2@o
/
-288'l
\/c;= 6 (/2)(18 4) (7o 7 ) + 2e8 .
/#C (s.s-Mi ve, = u +se.2=s,.s v
&m (zs-is) , dia7. en su t ve w) >
Ycw= b'cl(3.5Nf' + o.3 (pch .
(pc = Oy = /486 pai (comp) , d= 0 20 s ss-28 8'
( > /8. d")
Mcw = te (28.8.) (3.sTEooo + o.5ni488) vc- = 24 o *v, > vc
- . vc' = vci = .s 7.s cA
- lo$c.$ Capac;G e / Coisc + ferr) = 576 + St.2= 894*b espe t, e er.4 "/i > vk e se.s y .. dai ce . 5
~ ~
~
Supplement 1 8-10-76
- w. ,e y -,_. -
r, *--
w = supe =pr g . _,,#-d. f ."t4-de'""" " "
i
- i. SECTION 2.4
- 1. In the paragraph in the middle of Page 2-4, you indicated that for structural integrity test and accident condition load combinations, stresses for sustained loads cannot be combined with those due to rapidly applied loads internally in the program and are combined externally. Provide an example of actual design to show how the stresses are combined externally and illustrate the combination on a stress-strain diagram.
Answer: See Attachment 2.
- 2. On Page 2-5 under Item 'b Creep, it is indicated that aus a result of concrete creep there is a reduction in concrete stress and an increase in liner stress.
Since the liner is relatively thin and may buckle under prestress, the liner should not be considered to contribute any strength to the containment vessel.
However, in the design of the steel liner, strain dur to creep of concrete should be considered to check its leaktightness *n6egrity. Revise the concrete stresses in the report if they hav.~oeen reduced, n
Answer: A reduced modulus of elasticity of concrete has been used in the analysis and thus the effect of creep on concrete and liner stresses has been accounted for. Our analysis indicates that for the load combinations D+F and D+F o+ T the concrete stress is increased if the liner is removed in the analytical medel. From the standpoint of concrete stress behavior for the SIT and LOCA load combinations, to remove the liner from the analytical model is not conservative.
The figures in the report have been modified to provide a comparison of both results at selected points.
- 3. Provide the procedure which you used in the design of the steel liner. In Table 2-2, you stated that no criteria on liner strains were used in the original design. Indicate the criteria you used for the steel liner design.
Answer: Table 2.2 has been modified to reflect li ter design criteria.
S-4 Supplement 1 8-10-76
- . m--- ..~ ..-,x _ .m_m __ _ . . _ _ _ _ _ _ __
w
[ 4. . Discuss in detail the effects of creep, including the following consideration:
Because of the different level of prestress in the wall in the vertical direction, the wall in the hoop direction, in the ring girder and in the dome the EE is different in all these directions and this effect should be considered in the analysis. The wall acts as an orthotropic element.
The different parts of the structure have simultaneously different E{
due to different specific creep. '
Answer: The effect of creep has been accounted for by the use of reduced modulus. Although the different parts of the structure have different prestress, the specific creep (creep due to unit psi) should be the same for the same material. Thus the reduced modulus should be about the same for the various parts of the structure. A calculation is attached to demonstrate this (See Attachment 3).
- 5. In Table 2-3 add load combination equation for repairs. This equation should include the seismic load term.
Answer: The FSAR and the current ASME Code load combinations do not include earthquake effects in combination with construction loads.
f i
I e l Ii l
I l
. S-5 Supplement 1 8-10-76
- A 8 IT' " *
{g' g . gg
-#TM. g,* Y '**X*
_+ -
-YT # #
Shast 1 of 2 ATTACIMENT 2 FOR ANSWER TO QUESTION 2.4(1)
N '//tM V2 [ 0/=~ $/T O ?i !) i f}7lC.*/
2 A" COME 3lT l55f'
- i?,':s 4-l4 7Ni:t !7 S e = *;.',' J '.4 / 0 '1lc = 0 2o D+F Q = /.}.0 y /0 ' N s Uc. = 0 20 f./G P
.== p r ?O M /C 1 'L/.k : 03o 2
I
\
i
(_
Supplement 1 8-10-76
, mow; pev w- <
--- - - - - ~ " ' - ~ ~ ~ ~ ~ ' ~ ~
She t 2 of 2 ,
ATTACHMENT 2 FOR ANSWER TO QUESTION 2.4(1) (Cont'd)
H m, . m M1 3 .+
s
- g' '
.w :,
- g. , - !!2*
- m as .. . .
i I'y ,E Q.
se O y -
- + -728(nton)
..,\ \
=
!y},
l a i.- _ _ _ _
a -2sio
+ cm ,
+ rs -
7 *?
_.$ ; -27eod +i9 dss - 7771 W 'f D+F I. lSP S/T Ee =2 7y/c a Ee = 4 0 sto' b + f + /./GP E 4= Eo vio ' 2 4= 20 x /o '
.a I -/e7
. R
[-637 n
I
-82' '
D v
L eJ._.
A
+ , , -6EB(nlern) '
'e_L \
sf a
_ ' g' . _ . _ _ .- __, . . _ __ g ..._.._ _ ___ __ M I* ! CE 9.g
~2V)o +21sz I
- Ms8 ij k l
b//221~be'Gi*14 de,o -777/-o.z(-sssa) , -0. won 4 f' =
so nos bG= -Gd58-0 8 (-7721) , -0 000/38h =
30 y to '
I IF Cone. 9tairrs !
$0 9 ~'23/O ~0 2c(-scq) , f 2494-c.2 04) ,-O.000l95h~ !
2,7y/06 4 y /o' Se = - SC~'7 -0 2 h2Sio) ; l'1-0 Z (2496) _
0 000] 5 7l ^
! 2. ~} % /C' 4/ /06 Y1Il G /T 0/10. $o/d Gins Carir1Cl b d CQlc. fDn1totYcas;et"lNWS
' Supplement 1 8-10-76
~~
g .._,a _
~
t Shsat 1 of 3 ATTACHMENT 3 FOR ANSWER TO QUESTION 2.4(4) - CREEP Creep of the concrete under sustained loads D, Fv, FH, and Fp has several effects on tendon forces and containment stresses. The most obvious is to decrease the tendon forces with time. This effect is taken into account in the prestress loss calculations.
Another effect is to decrease concrete stresses and to increase liner stresses and strains, which are compressive over most of the containment structure.
The decrease in concrete stress is due to the additive effects of the decrease in tendon force plus the creep straining of the concrete acting with a non-creeping liner, which tends to shed compressive stresses from the concrete to the liner.
This latter effect is taken into account in the analysis through the use of the effective Young's Modulus, E[, appearing on page 4-3 of the report.
Using this approach, less concrete compression is calculated to be available to resist SIT or LOCA conditions than would be calculated by considering the reduced tendon force alone.
With respect to liner stresses and strains, the structural analyses show that the EE effect (E& = 2.7 x 106 @ present and El = 1.8 x 106 @ 40 yr versus Ec = 4 x 106 " instantaneous) is much greater than that of the reduced tendon force. The net result is liner stresses and strains which have compressive values much greater than those which occur at initial prestress. The liner strains in the report include this.
A third effect of concrete creep is to produce creep induced stresses which result only when the EE is not uniform over the containment structure. If EE is uniform, the stresses at any time are equal to those at initial prestress less tendon losses, in the case where the liner is not part of the model.
The E[ values used in the structural analyses correspond to specific creep values, se, which were calculated based on the 1) average age of the dome concrete at application of the dome prestress (average) and 2) the duration of this prestress to "present" and to "40 yrs". The resulting E[
values were applied to the entire structure in the analyses for D + F. It wasrecognizedthatadifferentE[isassociatedwithvertical(F), hoop y (FH), and dome (Fp) prestress loading conditions. This is so only because the concrete age at application of each prestress load is different and the duration of each type is different. However, Ej values were based on the dome concrete age and dome prestress since it is that part of the containment structure which is most effected by the delamination. Also, it was felt that these EE values would be an average for the wall since, chronologically, FD was applied between Fy and HF . Nevertheless, a more accurate determination of the' creep effects due to the separate application and duration of the nrestress is discussed below.
As pointed out previously the determination of EE de' pends on 1) age of concrete at loading and 2) duration of load. EE is independent of the level of stress in the concrete, which is reflected in se (p in/in per 1 psi of stress).
Supplement 1 6-10-76 wwn - es 1 - -- -
m- - -
'i
~
f i
Sh: t 2 cf 3 ATTACICIENT 3 FOR ANSTER TO QUESTION 2.4(4) - CREEP (Cont'd) i
/*
Therefore, creep induced stress under either y F , F , rFD will be reflected only in differences in E' for the various elements Hin the containment for i each of these prestress Soads. The total results would be obtained from the sum of the analyses shown below.
Fo l Fv u i F
Adv Adv Adv s- _y_._. -
A qv Agv g Argv l i
=
Awy -
- Awy V 7_. A
_ WV
. FH ,
Awv = age of wall at time F y is applied.
Argv = age of ring girder at time Fy is applied.
Adv = age of dome at time yF is applied.
Similar for AwH, ArgH, AdH and AwD, ArgD, add.
Dv = duration of Fy from time of application to "present" or "40 yr" times, similar for D Hand D D Knowing the values of A and D permit calculation of E' for the three elements.
l Supplement 1 8-10-76
-4f-Q I M' 'I*- * *O#'*#
She:t 3 of 3 ATTACHMENT 3 FOR ANSWER TO QUESTION-2.4(4) - CREEP (Cont'd)
'/ Values for E' were obtained based on average pour dates for the wall, ring girder, and Eome and average stressing dates for the three tendon systems.
This is presented below.
Wall Rina Girder Dome Average Pour Dates: 6-19-72 9-1-73 5-15-74 Vertical Hoop Dome Average Stressing Dates: 11-15-74 2-15-75 12-1-74 "Present" Time "40 yr" Time Ej (psi) x 100 Ej (psi) x 100 P/S Aw Arg Ad D Ring D -
Ring System (da) (da) (da) (da) Wall Wall Gir. Dome Jzrsl Gir. Dome Vertical 880 425 180 545 3.13 2.89 2.63 41.5 2.08 1.96 1.75 Hoop 970 515 270 455 3.17 3.03 2.86 41.2 2.13 2.04 1.89 c.
Dome 880 455 210 515 3.13 3.03 2.70* 41.4 2.08 2.04 1.80*
- used in structural analysis of containment.
i The differences between E' values for the wall, ring girder, and dome under a specific prestress conditEon is not enough to produce stresses significantly different from those reported.
Supplement 1 8-10-76 pyw>seyApp 'a -
. ^:7~ ,* * -
- +~w-.'e-^ ~ ~ ~
v , _ . . _ _ _
e
/ SECTION 3.1
- 1. Discuss the reliability of direct tensile tests performed on cores. Since in the structure the radial tensile stress occurs simultaneously with two orthogonal compressions or with two orthogonal tensions, a more thorough investigation is required.
- Answer: The direct tensile test was designed to identify the tensile capacity of the concrete in the structure in relation to its compressive strength.
It was not intended to define the property of the concrete in a state of triaxial stresses, since the actual state of stress at points of strees concentration in the delaninated dome cannot be accurately defined.
The effect of the tensile stress in combination with two orthogonal compressions is discussed in Section 3.3.3 of the report. .No further investigations are planned.
(
4,.
S-6 Supplement 1 8-10-76 Y'Yf
_f l fY f *'*' ' S b'S ? h '
~ '
I SECTION 3.3
- 1. In the list of factors which may have contributed to the delanination problem, i add: creep and stress concentrations (at tendons) inherent in this type of structure.
Answer: Creep.in the membrane direction would not increase the radial stress.
The effect of stress concentrations is discussed in section 3.3.2.
- 2. In Section 3.3.2 it is indicated that' by using SAP IV computer program and l the modal shown in Fig. 3-16, the effects of material properties on radial tension stresses are evaluated. Identify in the model:
3 (1) the steel elements, such as reinforcing steel, and tendon conduits, i
Answer: There is no element representing reinforcing steel or tendon conduit.
The effect of reinforcing steel is calculated as transformed concrete area and represented by effective Young's Modulus. Modeling of the t.ndon conduit is described in Section 3.3.2. -4 l i
(2) the manner in which the prestressing force is applied, indicating if the prestressing force component tangent to the done curvature.is considered.
[
-Answer: Prestressing force is applied on three middle layers of the model in both the radial and the tangential directions of the done.-
- 3. Provide the hand calculation which you made to'obtain the radial tension.
Answer: These calculations are included in attachment.
i s
i 4. In Section 3.3.4, transient thermal gradients may generate shear stresses, and should be considered in the analysis. Similar effect exists for localized thermal gradients.
-Answer: Since thermal restraint produces normal strain, but no shear strain, the thermal gradient causes shear stress; but only in the areas which are reinforced for shear (Chapter 14 of Reference 11).
4 f
- 5. The solution for stress concentrations as shown in Fig. 3-17 & 3-18 is
- incomplete. It should be noted that compression exists also in the direction parallel to the conduit (c1). This stress generates additional stress concentration in the plane (02; c3) orthogonal to the tendon, which should be added to the stresses shown in Fig. 3-18.
I Answer: Assuming this question addresses the effect of Poisson's ratio, this effect was considered and is discussed in Section 3.3.2.
l C
l' S-7 Supplement 1 8-10-76 1 YY (Y y f' -
p N FN~ *' 5 ' ' ' ~
[
- 6. When the effect of tendon conduits is analyzed, it should be noted that this effect is different when evaluated in the direction parallel to the tendon and orthogonal to the tendon. In the direction parallel to the tendon a 1/4" thick pipe (5"$) approximately replaced the removed concrete.
But in the direction perpendicular to the tendon, the pipe introduces a flexible link which modifies the average properties of the concrete section.
Answer: We have reviewed the effect of the conduit on stresses following a path parallel to the plane of the membrane and results are illustrated in the attached figure. The distribution shown in the attachment indicates that
.the effect will not be significant. However, the effect of conduits are conservatively represented by a concrete layer with equivalent Young's
, modulus calculated by the ratio of net concrete volume to gross concrete volume.
('
k
( .
S-8 Supplement 1 Revised: 12-10-76 mm3.pm.w= , -s ~ ~ - -
ATTACHMENT FOR ANSWER TO QUESTION 3.3(3)
. -. The radial tension are hand-calculated as follows:
i' s
h y y"eqr-nsps a (3
s M :d A ()
.3 Np s$
o Top Tendon R = 1343.9" 1
Middle Tendon R 2
= 1338.4" -
Bottom Tendon R 3
= 1332.8" Tendon force at 0.7 ultimate = 1633 tendon spacing 30" "
Top Tendon o y = 1633 x 1000 12.1 12.1
= 40.5 x = 13.6 psi 30 x 1343.9
- 36 36 1633 x 1000 17.6 30 x 1338.4 x 36 = 40.7 x 17.6 , 79,9 p,i Middle Tendon o =
2 36 f
i Bottom Tendon a 2 3" 0 332 8
- 3j = 40.8 x j = 26.3 psi The radial tension due to all three layers of tendon are superimposed as follows "N.9 -- I3.G -20.S Ia.9
-14.5 3,5 O.h x
('gg,g 40.S 4 2
?GS -43.5 t
' 1.~I
(
Supplement 1 8-10-76 l
vw;4.wixw.u rra :m e - - sr. . ~ ~ ~ - - ~ . . . ~ - - . , , . - - . . .
. t ATTAC1D{ENT FOR ANSWER TO QUESTION 3.3(6)
(,
d il d4 '9 _ggg 7
- =
,z ~
-1G"10 PSt -, V*) _
~
= = @ i
-gest 9 y y F !
45 P51 d
-tG98
-1707
( -177.2
-n+l
~*
-nes e -191G g
N Q -\%50 z.su a 2.1st"R L S rREss PROFILE PARALLat To MEMBRME Supplement 1 Revised: 12-10-76
..,,n , ,_, . . . . - . . _ . , . -. _ .. . .-.- - . - -. ..
. i SECTION 4.4
- 1. In Sections 4.4.1 and 4.4.2 you indicated that in order to consider the containment structure serviceab.}e for the two loading conditions the shear -
capacity of the tendon conduit would have to be considered. Such consideration may not be possible, unless the bond stress between the conduit and concrete can be justified to be adequate.
Answer: The tendon conduit is not required and has not been considered as
. contributing to the shear capacity.
/'
(
S-9 Supplement 1 l 8-10-76 i
ky - f- 'b<N* ^5 g b %
' i l
e _
~l SECTION 5.3
- 1. In releasing the prestressing force as a result of tendon detensioning, strain recovery will occur. However, most likely the strain recovery in concrete will be resisted by the steel reinforcing bars and steel liner, because of creep effects, and tension may result in the concrete. Provide an analysis to show chat the resulting cracking in dome concrete will not jeopardize the structural integrity of the dome particularly in the region of the liner anchors.
Answer: Not applicable under new repair sequence.
- 2. The behavior of the detensioned dome is strongly influenced by the creep of the prestressed structure which has taken place after prestressing and up to this date. The detensioning of the dome will not return the structure to a previously unprestressed state, whatever the sequence of operations.
It is ther efore imperative to analyze the detensioned dome for the influence of creep. Present such an analysis and demonstrate that the integrity of the detensioned dome will not be impaired. The analysis should include the -2 ring girder and the top of the cylindrical wall.
Arswer: Not applicable under new repair sequence.
(' 3. The figures 5-11 to 5-14 do not include a study on shears. Provide a detailed analysis of shear stresses in the detensioned dome and demonstrate that these shear stresses, acting simultaneously with normal stresses, do not endanger the stability of the dome. Special attention should be given to radial shears.
Answers: Not applicable under new repair sequence.
- 4. Either justify in detail the use of 24" for the dome thickness in the present analysis, or present a parametric study for different thicknesses; for instance 24"; 18"; 15".
Answer: The response of the structure to detensioning, and the parametric studies of section 3.0 indicate that the structure is responding as a 24" structure. The addition of epoxy grout, radial anchors and new reinforcing on the cap will assure its continued performance. Also see response to question 1.2.2 l
- 5. Demonstrate that the detensioned dome a;d eta el liner can take the load applied during the repair operatiot t .
l l
Answer: Not applicable under new repair sequence.
1 ,
i (
c I
S-10 Supp' pent 1 8-10-76
" ~ ~' ~*
- *e- j_ ,
( 6.
Present 2 detailed discussion of the provision made to monitor the behavior of the done, the ring girder, and the top part of the cylindrical wall during repair operations. Indicate:
- a. The acceptance criteria for safety in such operations, and Answer: This information has been added to the report as Section 5.0 Corrective Action.-
- b. the provisions made to safely stop the repair procedures if the acceptance criteria for safety are not met.
Answer: , All activities on the dome will. be temporarily suspended and no personnel, except inspectors, will be ' allowed on the dome after a stop work signal
- until approval to proceed is obtained from the Engineer and the Owner.
The acceptance criteria shall be in accordance with the requirements l noted for each measurement in Table 1. Work shall stop immediately when ,,
readings are outside tr. limits'noted in Table 1 for displacements and l liner strains and the Engineer shall be notified.
-An unsatisfactory set of readings requiring immediate notification of the Engineer during detensioning shall be when one concrete strain or
. reinforcing bar gage reading exceeds the values specified in Table 1.
The top surface of the dome shall be visually inspected for cracks before commencement of detensioning and any findings recorded. During detensioning and retensioning operations, the inspection for cracking shall be made on a daily basis as a minimum. Observations shall be reported to the Engineer.
- 7. Describe in detail the methods, acceptance criteria and methods of inspection for the grouting of the cap on the dome, the radial anchors to be. installed and the grouting of these anchors. Present the planned testing of these anchors.
Answer: Grouting of cap of dome is no longer part of the repair sequence. The
' procedure for sizing radial anchors is described in Section 5 of the report.
4
.A test program is being conducted to choose the best set of anchoring devices among the following; a cone and expansion shell anchor system grouted with cement' grout, a thread rod with nut bearing grouted with cement ' grout, . a threaded anchor with nut bearing grouted with epoxy grout, and a' deformed rod grouted with epoxy grout.
l \'
S-11 Supplement 1 8-10-76
$ 4mwmrw&wn mw .
= ~~
. ~ . , --
v
[,
Three (3) types of anchors manufactured by Williams Form Engineering Company have been selected for testing.
-1. Williams long cone and long expansion shell (LCS-200).
' 2. Williams standard cone and standard expansion shell (SCS-200).
- 3. Williams deformed anchor with and without an end nut. In addition
- a non-deformed anchor with nut bearing assembly will be tested.
Three~(3) different grouts are being tested:
- 1. Masterflow 814 cement grout.
2.- Masterflow 713 cement grout.
- 3. Sikadur Hi-Mod 370 epoxy.
Following series of tests are conducted to verify the anchor strength. a
- 1. A shallow hole, 2" in diameter and 7 inches deep.
- a. To verify that torquing of bolt will not cause damage or rupture to the nearby concrete.
- b. To establish'the failure mode of concrete for available minimum depth.
- c. To establish the design load capacity of the anchor at the minimum available embedment depth.
. 2. A 2" diameter hole 10 inches deep.
4
.a. To verify that torquing of bolt will not cause damage or rupture to the nearby concrete.
- b. To establish the failure mode of concrete for this embedment.
- - c. To establish the load capacity of anchor for this embedment.
- 3. A 31 inch deep hole.
- a. To develop and maintain the design preload in the bolt.
~ '
' - b. Upper and lower bound torque values requirements to develop the design preload.
7 k
l .?
( . N ..
L S-12 Supplement 1 8-10-76
- M r4 S'13PN6-r wg y q,.
Y *99' g[ 4fepi7 Mg pg ..
E ,
l _
9 _
.-~3- ,
- l l
(-
l
- c. To verify strength of the anchor with respect to concrete capacity. '
- 4. A 31 inch deep hole (epoxy grouted test block). l
. I To investigate anchors capacity in epoxy grouted concrete.
- b. To establish the failure mode of concrete.
- c. To compare the anchor capacity with solid concrete block.
The most suitable anchor type will be established af ter testing is complete.
Final anchor configuration and design basis for the anchor will be submitted as an addenda to the report.
- 8. Provide a commitment that sufficient strain instrumentation will be installed at the top and bottom of the dome to assure that during retensioning of tendons the upper portion of the dome (above the crack) will be participating in developing compressive stress at the same rate ,
as the lower portion.
Answer: The instrumentation is described in Section 5.0. The gages which exist in the cap will be replaced with strain gages on embedded
. reinforcing bars and the radial anchors. Observation of this f' instrumentation during the retensioning and SIT should assure that the structure is responding as designed.
- 9. Indicate in more detail the planned method of waterproofing of the repaired dome and its protection against detrimental environmental conditions.
Answer: A detailed description will be provided later.
- 10. Describe the acceptance testing of the repaired dome and the inservice monitoring of the structure.
Answer: Acceptance of the repaired dome will be based on satisfactory completion
( of the SIT. Af ter the SIT the currently accepted inservice inspection requirements will be performed.
- 11. Investigate the influence of possible cracking in the hoop direction on the dome tendon conduits.
Answer: Not applicable under new repair sequence.
\
S-13 Supplement 1 8-10-76 5 .h ' '
-- ~ * '
t-
.(~
TABLE 1 FOR ANSWER TO QUESTION 5.3(6b)
PREDICTED STRAINS'AND DISPLACEMENTS 15% PRESTRESS Gage Type of Predicted No. Measurement Measurement Range E22 Liner Rad. Displace. 0.017 in 0.004 in E23 Liner Rad. Displace. 0.017 in 20.004 in E24 Liner Rad. Displace. 0.017 in 10.004 in E25 Liner Rad. Displace. 0.008 in 10.002 in E26 Liner Rad. Displace. 0.008 in !0.002 in E27 Liner Rad. Displace. 0.008 in 10.002 in E28AV- Liner Vert. Displace. 0.041 in 10.01 in E28BV Liner Vert. Displace. 0.016 in 10.004 in E29AV. Liner Vert. Displace. 0.041 in 0.01 in E29BV Liner Vert. Displace. 0.016 in 0.004 in E30AV Liner Vert. Displace. 0.041 in !0.01 in E30BV Liner Vert. Displace. 0.016 in !0.004 in Apex Vert. Displace. 0.129 in 10.032 in 15' Radius Vert. Displace. 0.129 in 0.032 in 30' Radius Vert. Displace. 0.120 in 0.03 in 45' Radius Vert. Displace. 0.069 in 0.017 in R118M Liner Merid. Strain 65 y in/in 33 p in/in R118D Liner Diag. Strain - -
Rll8R Liner Hoop Strain 33 4 in/in 2 16 4 in/in R119M Liner Merid. Strain 65 p in/in 32 y in/in R119D Liner Diag. Strain - -
R119H Liner Hoop Strain 67 p in/in 1 33 p in/in R120M Liner Merid. Strain 65 y in/in 33 p in/in R120D Liner Diag. Strain - -
R120H Liner Hoop Strain 33 p in/in i 16 4 in/in R121M Liner Merid. Strain 65 p in/in 1 32 p in/in R121D Liner Diag. Strain - -
R121H Liner Hoop Strain 67 p in/in ! 33 p in/in R122M Liner Merid. Strain 73 4 in/in 2 37 p in/in R122D Liner Diag. Strain - -
R122H Liner Hoop Strain 74 y in/in i'37 p in/in R123M Liner Merid. Strain 65 p in/in 2 33 p in/in R123D Liner Diag. Strain - -
R123H Liner Hoop Strain 33 p in/in 16 u in/in R124M Liner Merid. Strain 73 p in/in 2 37 p in/in
, R124D Liner Diag. Strain - -
R124H Liner Hoop Strain 74 u in/in 37 p in/in R125M Liner Merid. Strain 65 p in/in 2 33 p in/in R125D Liner Diag. Strain -- -
R125H Liner Hoop Strain 33 p in/in t 16 4 in/in x
Supplement 1 8-10-76
= ' .
A f, NQh / #' "P * " - ' ~ *-
.**"***"~W -*-w+=wem=*e- ema*? w
, ,, ~
s.
SUPPLEMENT 2 t RESPONSES TO STRUCTURAL ENGINEERING BRANCH g j"g*- --
g !
COMMENTS AND REQUEST -
FOR i ADDITIONAL INFORMATION ON CRYSTAL RIVER UNIT NO. 3 REACTOR BUILDING DOME DELAMINATION INTERIM REPORT AND SUPPLEMENT NO.1
- 1.
GENERAL COMMENT
S.
In the report the applicant discussed all possible factors which could have caused the delamination of the dome. No single or overriding mechanism has been positively identified as the cause of the delamination. However, the following facts are significant.
- 1. The indication of a tension failure along the delaminated surface.
- 2. The complete fracture of the coarse aggregate on the delaminated surface.
' 3. Large variations in the strength values obtained from the direct tensile tests of the concrete.
- 4. The presence of cracks of various sizes and extents in the concrete below the delamination as indicated by core borings.
On the basis of these facts, the sequence of events that led to delamination could be surmised:
From the evidence indicated above, one could conclude that; (1) the characteristics of the dome concrete are such that it is crack-prone, and localized cracks may have existed even before the prestressing force was applied, and (2) the coarse aggregates are fragile, thus, instead of acting as crack arresters, they became the path of cracks.
With the existence of precracks and the presence of fragile coarse aggregates, the radial tension accumulated from all sources was so large that it overcame the very limited tensile strength of the concrete, resulting in the separation of the dome concrete.
F nn.,nnn: e , . -.
.T m. -
SUPPLEMENT-2
-PAGE 2.-
( It'has been found by various investigators that cracking of concrete under compression is slight for loads below 30-to 50 percent of the ultimate. ~ This is basically the reason why the allowable concrete compressive stress is limited tcr 45% of the ultimate. The cracks, if any, which initially may-have developed in the dome concrete as a result of prestressing are unstable. They increase in length'and width until either they eventually stabilize i or ultimate failure occurs.. The slow crack growth in concrete under sustained loading is most likely associated with creep.
The postulation of the delamination mechanism and the understanding of concrete crack initiation and propagation are essential for the establishment of the dome repair procedure and its evaluation. The following repair pro-cedure is being purssed by the applicant:
- 1. Holes will be core-drilled into the lower concrete;
- 2. Top delaminated concrete will be removed;
'3. Final inspection of 24" structure will be performed;
- 4. Lower level cracks will be grouted with epoxy;
- 5. Radial anchors will be set and the holes grouted;
- 6. New reinforcement and concrete will be added;
- 7. 18 tendons will be retensioned;*
. 8. Structural Integrity Test wil'1 be performed. l
- The 18. tendons will be partially retensioned as described i in Section 5.2.9, Page 5-6, September 22, 1976 revision to the report, " Reactor Building Dome Delamination." l 1
On the basis of the postulation of the delamination I
. mechanisms and understanding of concrete crack initiation and propagation as discussed above, the staff has reviewed ,
and evaluated the repair procedure. However,.before the l staff can finalize its evaluation, the applicant should ;
respond to the staff's concerns as indicated below: l l
l I
i 1
w e m p y wy m mm w n . r: m unr- m y. , , ---n , .n, - ,
, , - - , -,-. , , - . . , , ,---n -
(. . .,
SUPPLEMENT 2 PAGE 3.
7 s
II. DOME REPAIR.
- 1. An analysis of the repaired dome should be made for the following conditions:
(a) Before the hardening of the cap concrete.
Answer: Analysis of the repaired dome before the hardening of the cap concrete has been performed.
The controlling stresses and deformations are reported in Appendix G, " COMPARISON OF DESIGNS,"
.Pages G-7 through G-9, September 22, 1976 revision to Dome Delamination-report Refer to column headed " Dead Load Plus Prestress at Early Plant Life."
(b) After the hardening of the cap concrete, including all the loading conditions a, described in the FSAR.
Answer: Controlling analytical results for the repaired structure with the new cap in place are summarized in Appendix C, " COMPARISON OF DESTCNG," rages G-7 through G-9. Other FSAR load combinations have r
not been presented since they do not control any of the final dome design.
Indicate the stresses and strains in the mainly reinforced concrete cap portion and in the prestresse> concrete lower portion.
- Answer: Appendix G, " COMPARISON OF DESIGNS," includes the requested information.
- 2. Provide a description of the final design of the radial anchors and indicate how the combined action of the~ cap concrete and the lower dome concrete is-ensured.
Answer: The final design of the radial reinforcement and the combined action of the cap with lower dome concrete are presented in Section 5.2.7 (Page 5-5, September 22, 1976 Revision).
Specific reference is also made to figures 5-22 and 5-23, as well as Appendix I.
-.m t#A .,
,,F '< .~,:'sr"-
. - " * " ~ ~ ~ ~ ~
(,
r SUPPLEMENT 2
.PAGE 4. *
,.f
-3. It was indicated that two layers of reinforcing steel will be provided in the cap. For the meridional reinforcing steel, if only one layer can be spliced to the existing meridional steel near the ring girder, indicate how the other layer can effectively carry the load if it is not spliced to the existing steel, noting that under internal pressure, dome concrete may crack -
in tension.
Answer: The #8 lower layer meridional reinforcement '
is provided for crack control only.
Figure 5-20 illustrates meridional steel provided versus that required and does not include consideration of the #8 lower layer meridional steel shown in Figure 5-19.
The lower layer of the meridional steel therefore is not assumed .o "... effectively carry the load...". The top layer of meridional and both layers of hoop reinforce-ment in'the new cap are considered to provide strength.
e 4. Since the repaired dome becomes a unique structural element of the ' containment structure, indicate any special con-siderations to meet the requirements of Regulatory Guide 1.18 in executing the structural integrity test of the containment.
Answer: Regulatory Guide 1.18 requires that displacement be measured at the apex and spring line of a containment dome. The instrumentation for the Crystal River Unit 3 Reactor Building has been considerably enhanced with regard to the dome.
Refer to Section 5.2.1.c.
for detail on the dome instrumentation for the SIT. The additional measurements of dome displacement will be in-cluded in the SIT acceptance requirements. The predicted response data was supplied by letter of October 8, 1976 (Attachment 1).
- 5. The original dome design concrete strength, f'c is based i
on 5000 psi; now a concrete strength of 6000 psi is used for evaluating the repaired dome. The basis for using 6000 psi is that the actual strength of the existing
- structure possesses-that strength. It is a well-known fact that concrete strength increases with age beyond 2S days and stabilizes after a certain time. Generally, designers of
~
l concrete structures do not take such increases into considera-tion construction.
and mainly to offset " ignorance factors" in areas of design r -
Revioed: 12-14 76
...7-7 Q- 7
--. -. ' v a' ~ ~ ~ ~b'%[
p SUPPLEMENT 2 PAGE 5.
r 4
Provide a justification that- such additional' margins of safety are not required in.the case of a concrete containment, noting-that there is a reduction in dome concrete area due to the presence of cracks, sheathing ducts and other possible voids, and if such reduction of concrete area is disregarded in the stress computation, the computed membrane compressive stress.may be less than the actual.
Answer:- The in-place concrete strength is usually not taken into account in design of structural concrete. The reaso'n for this practice is that the in-place strength is not known at the time the design is performed.
However, it is also current practice to use a design strength (f'c) based on an age closer to the time of first service loads rather than based on an arbitrary age (e.g. , 28 days) .
For the Crystal River Unit 3 Reactor Building Dome, the in-place strength has been evaluated in accordance with the accepted practice of Chapter 4, Section 4.3.3 of ACI 318-71 and the compressive strength has been determined to be 6130 psi (See Table 3-2, Page 3-15,
' Dome Delamination Report.). Another calculation, using ACI 214 (Midcell Method) and ACI 318, Section 4.3.5.1, had given a compression strength of 6600 psi (See Page C-5, Dome Delamination Report.).
Therefore, there is sound technical basis for using a design in-place compressive strength of 6000 psi.
!?ith regard to ". . . presence of cracks , sheathing ducts and other possible voids...":
1.- The. lower level cracks are parallel to the membrane and do not constitute a reduction in the concrete area available to carry membrane forces. They have been successfully grouted (see Attachment 2).
- 2. " Sheathing ducts' are 5" diameter Schedule 40 pipe and replace the displaced concrete. See Supplement 1, August 10, 1976 revision, page 5-8, Question 6'for additional detail.
- 3. We are not aware of "other possible voids." Con-sidering the number of cores taken in the Crystal River Unit 3 dome (in excess of 2000), it is unlikely that any voids exist in the' dome' .
W e' j. ' , p YNh .. _
I *#-
!4.. b .
p .
t
-SUPPLEMENT 2 PAGE 6.
-(
Considering the above 3 factors and the actual response of the structure to 15% detensioning program, " computed membrane compressive stress" should be quite close to actual stress seen by the structure undcrany load combina-tion.
- 6. The cracks in the dome concrete as discussed in the general comments have reached stability. The Structural Integrity Test (SIT) will affect such stability. Provide an evaluation of SIT on the lower level cracks of concrete which may not be grouted with epoxy. Provide the data on the effectiveness of epoxy grout in controlling concrete cracks.
Answer: The current through-thickness stressos in the dome are compressive (see Figure 5-22, September 22, 1976 revision). The pressurization of the Reactor Build-ing for the SIT will increase the existing radial compression through the entire thickness of the repaired dome. The added radial compression will vary from 63.3 psi on the inside surface to zero (0) on the outside surface. Since the through-thickness stresses wi'l still be compressive, they will not disturb the stability of the lower level cracks.
Although not essential to the structural behavior during the SIT, the epoxy grouting of lower level cracks has been accomplished (see response to Item II.5) and should enhance through-thickness stability.
III. CAUSES OF DELAMINATION.
- 1. On Page C-3 in Appendi. C under the subsection on " Direct Tensile Test Results" the applicant indicates that the range of direct tensile tests on 6 core samples was 230 psi to 505 psi with an average value of 420 psi. In view of these low results, the allowable membrane tensile stresses in-dicated in Table 2-2 appear high. Discuss the cause of these low tensile ultimate stresses, the reason for the wide scattering of the test results and the possibility that the delamination phenomena was caused by the poor quality of the aggregate, and the propagation of local cracks along the whole surface of the dome as surmised in the general comments above.
l l
,,w y _,- ,,n ...,.n w < .-v~n u- - - e ~~~ ~~ ~ - -
- .~ - - . - .. . - _ . -
~~
- (
SUPPLEMENT 2 PAGE 7.
( ,
Answer: The variation of direct tensile test results is dis-cussed in Appendix C of the report, " Reactor Building Dome Delamination." The Table on Page C-15 (Attachment D) presents direct tensile strength test results and des-cribes in the remarks column the relative " hardness" of the coarse aggregate. A review of that table indicates.
tensile strength is related to " hardness" of coarse aggregate. Also, see Page C-6 for a discussion of direct
' tensile tests by Mr. Joseph F. Artuso.
With regard to the tensile load capability of the concrete, two types of tests were performed to measure the tensile capability of the "in-place" concrete; i.e., split tensile and direct tensile tests. Attachment B of Appendix C of the Dome Delamination Report indicated that the average value for split tensile test of the "in-place" concrete was 710 psi, with a minimum of 625 psi. Attachment C of Appendix C in the Report indicates that the average value for direct tensile tests of the "in-place" concrete was 420 psi, with two test values lower than the average; i.e., 360 and 230 psi.
,e As indicated on Table 2-2 of the Dome Delamination Report, in the original design criteria the allowable membrane tension stress for factored" loads was 212 psi and zero for service loads. As indicated above, the lowest individual value for tensi.le strength obtained from either the split tensile or direct tensile tests was greater than the original design values for membrane tension for even the factored load condition.
The. quality of the aggregate and the propagation of local cracks along the whole surface of.the dome has been dis-cussed in Sections 3.3.lc and 3.4 and Appendix F of the Reactor Building Dome Delamination Report as being a contributor to the delamination. However, Chapter 3 of the report discusses several additional factors which may have contributed to the delaminated condition of the dome.
i l
Revised: 12-10-76 4 ,
a gur DF
..._._,".,"t.*'" ,.
n SUPPLEMENT 2 PAGE 8.
/-
I The~ dome (pours J through Q) did not contain radial reinforcement which cracking.would have prevented gross propagation of laminar Radial ties have been incorporated into the repaired dome to resist predicted radial stresses (see Section 5.2.7).
2.
The applicant presented in Fig. 3-22 the plane strain finite element trations at themodel tendon used to evaluate some stress concen-ducts.
- a. Present a detailed description of boundary conditions .
(especially at the duct) and initial conditions introduced in the computer analysis for all cases of stress con-centration.
Answer: The model shown in fig. 3-22 was used to calculate stresses in the concrete due to shrinkage effects.
At the interface of concrete and duct, perfect bond was assumed because of compressive interface The outside boundary was assumed to be free. pressure. Rollers on the boundaries were used to simulate symmetry.
The model was assumed to be stress-free prior to
, 7- . application of the shrinkage effects. The geometry and material behavior was assumed to be linear.
- b. Justify the use of plane strain to analyze what is essentially a three-dimensional problem.
Aaswer:
i The plane strain model is not intended to accurately describe of conduit, the real situation (for example, 3 layers double curvature and loads induced by the tendon in the conduit). It was, however, con-sidered adequate to examine the replacement effect of the 5" Schedule 40 pipe.
Revised: 12-14-76 g qy..g .6 W'6 '
, ,e - s
' .O INSTRUMENT LOCATION
.r The location of displacement measurements are as follows:
x
_ Cylinder Wall and Dome Junction Radial Displacements
. LP Gage Loc. Elevat. ' en, Azinuth Note, 1, 2, 3 98'-O' 90 , 2000, 333 0-55' Radial Displacement 4,5,6 108'-0" 90 0, 200 , 333 -55' Radial Displacement 7,8,9 140'-0" 90 , 2000, 333 -55' Radial Displacement 10, 11, 12 172'-0" 90 , 200 , 333 -55' Radial Displacement 13, 14, 15 204'-0" 900 , 200 , 333 -55' Radial Displacement 16, 17, 18 236'-0" 90 , 200 , 333 -55' Radial Displacement 19, 20, 21 246'-0" 90 , 200 , 333 -55' Radial Displacement 22, 23, 24 253'-0" 900 , 200 , 333 -55' Radial Displacement 25, 26, 27 267'-0" 90 0 , 200 , 333 -55' Radial Displacement 128, 129, 130 270'-8" 90 0, 200 0, 333 -55' Radial Displacement Ring Girder Vertical Displacement - LP LP Gage Loc. Elevation Azimuth Notes 28, 29, 30 267'-6" 900, 200 , 333 -55' Vertical Displacement Dome Vertical Displacement - LP LP Cage Loc. Elevation Location Notes 34_ _282'-4 1/8" Dome Apex Vertical Displacement 128, 129, 130 49'-.3" radius 900, 200 , 333 -55' Vertical Displacement -
164, 165, 166~ 28'-8" radius 90 0 , 200 , 333 -55' Vertical Displacement Equipment Access Opening Displacement - LP
'LVDT Gage Loc. Elevation Notes 35, 37, 38, 39
- 40, 41, 42 132'-0" Radial Displacement s
35, 37, 38, 39, 40,;41,-42 132'-0" '
Vertical Displacement
- 1-
,,, s.m y m~ -~ s- - - - -
-- h -- " ~
(
^
2 P Jtquipment Accean Op ning Displacement - LP (Cont.)
LVDT Cage Loc. Elevation
/
Notes x 36 120'-0" Radial displacement 36 120'-0" Vertical Displacement 43 144'-0" Radial Displacement 43 144' -O" Vertical Disp).acement 44 147'-3" Radial Displacement 44 147'-3" Vertical Displacement 45 151'-6" Radial Displacement 45 151?-6" Vertical Displacement 46 155'-6" Radial Displacement 46 155'-6" Vertical Displacement 47 159'-6" Radial Displacement 47 159'-6" Vertical Displacement 48 163'-6" Radial Displacement 48' 163'-6" Vertical Displacement e
e a
v m.m c_. . . . . . ,
c, m - -
,. ,e ,
,-~ . .
Displace = ant Acceptance Criteria ~
(
.' Cage Theoretical Limiting .
Measurement Displacement Displacacent ' Tolerance ID (inches) (inches) (inches) 1,2,3 Radial 0.010
, ~ 0.020 0.010 .
4,5,6 Radial
~ - 0.090 - 0.115 0.025. - .
7 , 8 ,~ 9 y.- ..
Radial. 0.205 0.260~ ~
0.055 --
10, 11, 12 . Radial, 0.200 :0.250 -
0.050 13, 14, 15 Radial.
0.205 0.260 0.055 16, 17, 18- -
Radial
,0.160 0.205 ~
0.045~ .
19, 20, 21 Radial 0.060 '
O.08 0 0.020 *
'22, 23, 24 Radial
-0.025 -
-0.040 ~
0.015 -
25,'26, 27 Radial -0.055 -0.07d 0.015 28, 29, 30 Vertical 0.215 0.275 '
~
0.060 34 , Vertical 0.905 '
1.135 0.230 .
35 Radial 0.10 O.130 O.030 35 0.10 Vertical:
- 0.130 0.030 <
36 Radial' !
O.08 0.105 d.025.' .s. !
36
. ' - \
Vertical 0.10 0.130 . 1 0.030 37 Radini 0.12 '
'O.155 . 0.035 37 Vertical 0.10 0.130 0.030 -
38 Radial 0.115 0.150 - O.035 .
38 Vertical 0.10 0.130 0.030 39' Radial 0.115 0.150 0.035 39 Verti, cal 0.10 0.130 0.030 O
a l
... ._.. .. .. _. ... q . _ _ m t-
. ,, ,%,pa *.-y =9, , wp.v..*.y * ' '- . . _ . . -
,_..___',m,
. ,. - 7 ..-. .
-=.=----~~-: - ---- T
_ 4 .. ,
~ ,~ 1 t . . ., .
p,lgtncecent Acceptmice Criterts (centinund)
. ~ '
Gage Theoretical Limiting Measurement Displace =ent Displacacent-ID . Tolerance (inches) (inches) (inches)'
. . c 9- S 40 lbnh1. -0.11 0.140 0.030 t
40 ' . Vertical . 0.10
... , 0.130 0.030 .
41 Ra' dial O.11
- 0.140 ... 0.030 41 . Vertical' O.10 0.130 0.030 42 -
Radial 0.11'
- O.140 -
0.030 42 .V. ertical . 0.10
. 0.130 0.030 43 Radial 0.090 '.120 O ' -
0.030 -
43< Vertical . '
- 0.103 0.135' O.030 .
44 Radial ..
0.090 0.120 0.030 . -
44 Vertical 0.105
- 0.135 , 0.030 -
45 Radial 0.095
^
.0.125 , 0.030,
~
45 Vertical 0.110 0.140 . 0.030 -
46 Radial 0.100
- 0.130 . 0.030 -
- 46 Vertical 0.11 -
0.140 0.030 47 Radial , 0.105 0.135 ~ 0 030 47 Vertical 0.115 0.150 5.035 ' .
~
48 Radial 0.110- 0.140 0.030,-
48 Vertical 0.120
~
O.155 0.035 128, 129, 130 Radial 0.015 -
0.025 . 0.010 123, 129, I ~10 Ve:cIcal 0.365 ' O.u0 0.095 164, 165, 166 Vertical 0.900 1.130 0.230 '
M A,
9*
. ... .. ~.. .. -. . . . . . .. , .. __.,._ . ..
., _ - .ppmm -%-- - ~ - ~ ~ ~ ~ ' ' ' ~ ~
. ,4 3 '
_ _ = . ~ . . a. 4,a e
-t Ms { g.m. -l FY An' en e:r.h cn*=sdateei ' ^' .t eWsd**** s-*i=* :. 4 5 T '.**-tw-st.s.t s 1
. . . M .. . .
ATTACHMENT 2 gQ, , .a- .
- r. :. . . i. . . (UMN 3htW Mttn;* ~. . .*
r, e,s m.--
7,.in. +. *
. . . . , ~ . .
. , 1,
],, -
7p - ' ~J'.
10h2525esce/ Oc,cLwl, ow . tun -
k (S
6 x . ii. 9 - -
g4} m-er/ '-
A .
-x septemberg3J1976v. -
1- f
. W G~o.
' i L2 y.
c.7,-
- p .-.. , ... .. ..
pr- 4..
. . ,,p .; ,.. - c -
~
e
. -g 3 4
.M- 'Dr;*.F/ TJ.Nrerw11thv . ' 7".. -
51f* .,a :.- 9 ~C11bertf AssociatespIrtti.wNP . .
-" +% *-# p.L7W '- *V** N - U{ - -
j.
i-igg .:?;. -wpny;op not:1498 mw- 'e ' ~f - "-
W.v. N.-x THeading",tFa.2,396039.e+1W M'? 5 '."+~ - m~ r e -r'3g.p .q' gy+e 'j_ .!i.
, , b . t':
.,.,s.;2.m .. y
. ..;-:~,+.. w : . - .
,.... . y: .
- 4. - -
. ~. t f.
hydk w?%3ahjectzW CR.7 D&mrB@al@mwa g
t-a W.e.fa.
y w +.y:,m.- -?
W- .R ( qi
,.~f ..-Epaxy Grouttag-u ~. ; .? ~ rW-a ~ e. k -- ,-. n >.- ~. w - y y q @.q.
q.yhy .w r,g,wDec. 7.h.9.trai endithuee"-M 7 -"-!<.wrm.w .7,.w.m.va.r-mw s ' m -r ~ - W --- ' ~
f.
eiza mA E.
..w.~.n.s.m W .w.e +.v.4.v m w w u'e.=u.v n.. w &i }c i@
ha.1 t,J,3;.mne.~x-c.w eu+ . yg.my..op1*rthei 4.poxycxrotding or Aclamitst ion e~:'+W-e-^
t s .n. % g . Q Q M e g . g g 7.g 1,y-she.:cy3 m ,e..w s.beem ~.~ 's" .m m.wy&
3-% 1.' i,?.'ta compIste success?. I:estimte thitte between 90:and J. ~ ' "
"' 5 y
'7d"W.I" .,'i 200't nr, the traeks are 'r23 led .with' very. strong eposy.?-- .
pp
. 'g_w t.# . c. . .:.* >- .. borded.t.oi hoth '. surfaces of the::c, racks;:.- ,
zy
.s
. . .. ....m..m. t y, x .. . ....v.. , . . . . . . .
. . . . v... a....., . . . . ,m.
- h9 J .
Y Dn ]0.$,IDNC M tfD W. 0;,,n.c g. $ hw. .
- . a 8' OW - 00 Y 5'c
~ D'3a d Mh.lar .. fhlsothes.ntireceerationt
,; o e actun*LIf. missed.(2):very.rew.;:ot.quite tr@resumbly.121ssed (for.r.. a.nu=ber:. .- . ,< .c .-
. . . . !; r
.... y f.l ai, .n. . McYord6gs.:show$og that holis hrul:hccu grouted}. s .: .
M+.Wp: ~t. .tooYany.grotat."and.{3) a nMcW3c1ccLlon or 10.4:u. .,.-: .. s- v M h
, ., w 's ,, y 0;
.y ..,
.,0 LhoIcs tbst.4md been .grooted and. redrilled accarta+ no .- % - .'. . . 4 e h, Lt.o .e a. .prout .whatever. at.-200cksic..c.h. 4.u. k.v..a ..,.. y. . .y g,ca._ .-M E
- .: g .. .n. : .. b i w. ;- - . 2
&L.u- n .-. . :n ?. n. : .
- r. -u:-..: .- ~ ,
~m. A t- Q 4.i h q eGJ. Details or disymbsermticassare givewin. Fiel:14 _,. --S..q-4u& e.n:
[ ca
{:.
w J
g4g1..;. base.beericaentMyouM:W.Malni %nmum; 4-# .,.~ ,y ,.m q j
- Qi.; hd .%wyW ? .%.~ . .m tr.a n::.M :" ?. ' $ - - : *: u - .~.
1-:-% m .%:27-- S%:.;5 l.h .
d.,t. .di
- . .n. e,d
<. I-r;~.4t M..:4mM..
O%. Un-C-+%:my'9.: ., , -..,.- . 953l,ghtrel,N.pg;c@c.,:
- - m.t.,.7.gaqQf- l{j.'
. . : c
.- -j,-2.a's ...;-;*&;,-:4.y:T ~-
- i ? -
., . l- } W. %i. ' -)%*tp5lfMe {M*.f*C4 $ . ;.-'- ~ .I
~
., v. u"e
- gfir.\,iar e.$.+d. & .-~.? ~s--+ *... ; , 9. .' .W.D:s I' . *
- %.-. s -
.$y' .- .W.-:..,v.." .
>. m y. ; -g ? %
%$eiebJ > e .i,. v-J.*i.4 .
s,- - -
9 t-W -
4
.,h?--W, -
- 6. . ; . , ' i:. *
'. ~< . .* * ,, e h .. .+-a- . , . - ;
v..ep.- w. -
1
- c. _-. . . . . . .,.
.J o h n C K D '. m
, - . . . . . .-_- 4 u.
yg. . :. . . . . - ?> - .. . . ,
.: c - .
s... . . . . . +
- /. ,e - _ y..
.*- - S-
,_- d, ,5 gr:
C..- 4
.* . . E
't
.. . m 4 r- ge W ] $55 M N'Nf M W Y s Sb ?-- $ see b
! VIBV 113 i 1
M UL i e.
- m.tno .yem.m n - a >~T-*".*^ ' " ~ ^ " ' ~ ' ' " ~ ' ' '
t
- - - - --