ML20207K446
| ML20207K446 | |
| Person / Time | |
|---|---|
| Site: | Fort Saint Vrain  |
| Issue date: | 12/12/1986 |
| From: | Connors G, Antoinette Lewis, Ryder R GENERAL ATOMICS (FORMERLY GA TECHNOLOGIES, INC./GENER |
| To: | |
| Shared Package | |
| ML20207K390 | List: |
| References | |
| 909190, 909190-IA, TAC-63576, NUDOCS 8701090423 | |
| Download: ML20207K446 (99) | |
Text
{{#Wiki_filter:Roll 2084 GA 1485 IREV.10/82) ISSUE
SUMMARY
TITLE O R&D 2 EFFECT OF FIREWATER C00LDOWN USING APPROVAL LEVEL REHEATER ON STEAM GENERATOR STRUCTURAL ODV&S T NT1:' "R T w G DESIGN DISCIPLINE SYSTEM 00C. TYPE PROJECT l DOCUMENT NO. ISSUE N0/LTR. 1900 909190 a M 22 CFL l QUALITY ASSURANCE LEVEL SAFETY CLASSIFICATION SEISMIC CATEGORY ELECTRICAL CLASSIFICATION
- Q1 FSV-I FSV-I NA i
g APPROVAL ISSUE PREP ^" ISSUE DATE DESCRIPTION / BY FUNDING APPLICABLE ENGINEERING QA PROJECT CW8S NO. PROJECT N/C NOV 0 71986 -- M- Initial Release
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NEXTINDENTURED Report Pages 1 thru 58 = 58 DOCUMENTS Appendix A Al thru A16=16 Appendix B B1 thru B24 = 24 Total Pages 93 N6757 I f 861230 F [ h 05000267 PDR ! REV l l SH l l
'EV I II l l l
__ l l ! SH 29 30 l 31 32 l 33 l 34 35 l 36 37 38 39 40 44 41 l 42 43 45 l 46 47 48 49 50 51 52 53 l 54 l 55 56 REV l l l l SH 1 2l3 4 5 6 7 8 9 10 11 12 13 14 15 16 17l18l19 20 21 22 23 24 25 26 27 28 [TROYERM.AAWOR] 5,6 PAGE 1 0F 98 l
909190'A i l CONTENTS l
- 1.
SUMMARY
. . . . . . . . . . . . . . . . . . . . . . . . . . . 4
- 2. INTRODUCTION . . . . . . . . . . . . . . . . . . . . . . . . 4 3 METHODS AND ASSUMPTIONS . . . . . . . . . . . . . . . . . . . 5 4
CREEP BUCKLING OF STEAM GENERATOR TUBES . . . . . . . . . . . 6
- 5. REHEATER HELICAL BUNDLE . . . . . . . . . . . . . . . . . . . 31
- 6. THERMAL SHOCK OF REHEATER TUBE SENDS . . . . . . . . . . . . 50
- 7. CONCLUSIONS . . . . . . . . . . . . . . . . . . . . . . . . . 54
- 8. REFERENCES . . . . . . . . . . . . . . . . . . . . . . . . . 55
- 9. CALCULATION REVIEW REPORT . . . . . . . . . . . . . . . . . . 56 APPENDIX A. THERMAL DATA . . . . . . . . . . . . . . . . . . . . A-1 APPENDIX B. MATERIAL DATA . . . . . . . . . . . . . . . . . . . . B-1 TABLES 4.1 Buckle Parametera "or EES Tubea . . . . . . . . . . . . . . . 11 4.2 Propertiea Used to Calculate the Creep Buckling Collapse Times "cr the Evaporation Tube . . . . . . . . . . . . . . . 22 4.3 Time to Creep Buckling Collapse (%) . . . . . . . . . . . . . 23 4.4 T;te Ovality to Cause Instantaneous Buckling Collapae (%) . . 25 5.1 Radial Thermal Displacement of Headers at Helix 1 Nozzle = . . 39 5.2 Reheater Support Plate Temperature . . . . . . . . . . . . . 40 5.3 Di= placement of Support structure . . . . . . . . . . . . . . 43 5.4 MODS AP Diaplacement Summary . . . . . . . . . . . . . . . . . 46 5.5 Reneater Tube Temperature at 50% Power . . . . . . . . . . . 47 6.1 Tabe Wall Temperature Determination . . . . . . . . . . . . . 53 FIGURES 4.1 Steam Ganeratcr Layout . . . . . . . . . . . . . . . . . . . 7 4.2 Creep Subroutine "or A800 Grade 1 and T2 Materials at Hign Temperaturea . . . . . . . . . . . . . . . . . . . . . . 9 Pa ge 2
909190/A l l l 4.3 Creep Subroutine for Analysis of Steam Generator at Very , High Temperatures . . . . . . . . . . . . . . . . . . . . .. 10 4.4 Steam Generator Tube Creep Buckling Times . . . . . . . . . . 12 4 4.5. Buckling Ovality Versus Time to Creep Collapse . . . . . .. 19 4.6 Helium Pressure History . . . . . . . . . . . . . . . . . .. 20 4.7 Helium Temperature History . . . . . . . . . . . . . . . .. 21
- 4.8 T2 Buckling Ovality Versus Creep Collapse Time . . . . . . . 26 4 5.1 Reheater Tube Model Input to MODSAP . . . . . . . . . . . . . 34
; 5.2 Reheater Support Plate Dimensions . . . . . . . . . . . . .. 37 5.3 Reheater Helical Tube Model . . . . . . . . . . . . . . . .. 38 5.4 Plan Diagram of Reheater Support Structure . . . . . . . .. 42 i
i I i 1 i 1 I i a i l j Page 3 l i
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l 909' 90 'A t I 1.
SUMMARY
Thia document preaent? the resulta o# a 9tructural evaluation o' the most critical regions of the steam generator during a single cycle cooldown from power using firewater in one reheater module. Simple but conservative methods were used to show that thi? one-time event will not violate the integrity o' the steam generator presaure boundary, provided that the local helium temperature ia limited to 1350*F maximum and mast not remain above 1300aF for more than 1 h,
- 2. INTRODUCTICN This document la the structural evaluation of the ateam generator during a single cooldown from power uaing firewater in one reheater module. The structural evaluation divided the event into two parta:
(1) the initial thermal shock o# the cold firewater contacting the hot reheater tubes, and (2) the subsequent heat extraction phaae when cool flowing water and hot helium generate temperature differencea between t steam generator componenta. The critical pressure loading caae is when helium preaaure and temperature combine with zero preaaure in the tubea. Thia load combinaticn may lead to creep buckling collap?e o# the ateam generator tubea. This case arisea if it is postulated that the ateam generatcra are vented to atmoapheric pre 9aure and the tube 9 are loaded by a varying external helium preasure. Thia external preasure, combined with hot modale and local helium hot streaks impinging on the tabea, requirea that a creep buckling =cde of " allure o' the tubes be evaluated. Thia is the only " Primary Stresa" loading identi*ied fcr this event. This one-time cooldown event la eqJivalent to an ASME Level "D" event and, as 92ch, thermally induced Streaaea need not normally be conaidered, unleas there la jJati#ication to do ao. The thermal porticn o' thia evaluation conaidered the temperaturea that occar 'rcm a cool-Page 4
l 909193 'A down from a power level that ensures that the maximum ateam generatcr tube metal temperature la less than 1350*F. The structural evaluations of thermally induced stresses have been deemed to be desirable beca Jse of the possibility of low ductility of the high alloy material in the reheater. Specific phenomena of concern are the aur* ace strains caused by thermal ahock in the cold-worked material of the reheater tJbe at cold-worked, small-radiua bends, and the plaatic strain accumulation in the helical tube bundle due to local yielding and elastic follow-up o' the adjacent tubing. The temperatares used in the evaluation were taken
- rom Re". 1, which gives helium and water temperaturea, and from Appendix A o# thia docJment which provides temperatures o* Selected structural membera during the event.
The purpose o# the evaluation la to show that the structural integrity of the primary pressure boundary of the steam generator will remain intact when using firewater in one reheater module to cool down the reactor. Three types o# structural evaluation were performed: (1) determination o# limiting temperature to ensure that creep backling o' externally pressurized steam generatcr tabe* doea not occur, (2) determination of local atraina in helical reheater tabea, and (3) determination o# the thermal ahock o' cold-worked, amall-radiJa bends in reheater tabes. 3 METHODS ~AND ASSUMPTIONS The methods Jaed 'or thia evallation are based on proven method 9 and techniquea Jsing, wnere poaaible, exiating "inite element comp 2ter modela or the ateam generator. Detailed analyais has only been per#ormed wnere abaolately necesaary. It haa been determined in Rer. I that #irewater " low will only be eatabliahed in one reheater mod)19 daring the event. Thia modlle waa Page 5
l 4 i 309130/A i l . l l evaluated for two loading *, thermal Shock loading and loada due to
- di'*eaential thermal expansion. The steam generator module =, which do 1
not have coolant in the reheater, will be subjected to the high helium 4
- temperature, including hot module and local hot streaks from the core.
i The uncooled steam generators were evaluated for high-temperature creep effects, specifically, tube creep buckling from an external helium i pressure on the tubes. 4 i The Sanicro 31 reheater tube material is a European Alloy 800-type j material which met the required material specifications of Alloy 800 $ Grade 2 at the time the steam generators were built. Properties from Ref. 2 have been used 'or the Sanicro 31 material. 4 CREEP BUCKLING OF STEAM GENERATOR TUBES l b The steam generator modules which do not have water flowing in the reheater will experience temperatures equal to the hottest region of the .i core outlet, mitigated only by the amount of mixing the helium encoun-ters between the core outlet and the steam generator tube being consid- { ered. The steam generator materials will be subjected to creep at theae j high helium temperatures; and, as the steam generator tubes will be externally preasurized, a creep buckling collapse failure mode may be { po*sible. To evaluate this condition, the GA BUCKLE computer code i (Ref. 3), was used. Thia code is derived from the PAN method (Ref. 4)
, which addresses creep buckling of long tubes under external pressure.
1 Each of the "ive types of steam generator tubing material
- were evalu-ated (see Fig. 4.1). The creep rate information and other material properties were supplied in Ref. 5 (included a* Appendix B of tnia document). Nominal tube diametera and wall thickne**es were combined with the worat-case tube ovality allowed in a small-radius, cold-worked tube bend, aa *peci#ied in Re e. 6. This la conservative, *ince tube backling will not occur in the
- mall-radiua bend because the double curvature of tube wall providea extra stif'ne**. The moat likely loca-tion *cr a tube to experience a buckling failure mode is adjacent to a 1
amall-radi2= bend. Thia la where the ovalization that occara at the ' Page 6
909190 ' A v HELIL:t 3 I :LET G h(, 3 (/r ::% ! REHEATER M I I M
~'
ALLOY 800 ! Fr- (SAN!ORO 31) I [ ' ]
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ECONOMlZ ER Q 1/2cr-1/2do steel
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r ,. y > HELIUBI Oi!TLET U Q.j. lp, (typical) Carbon Steel g# w; water Inlet a o o I ,- _ _ i g g } I, Main Steam
!<ainSteam$33tg7 Inlet Outlet StJan Fig. 4.1. Steam Generator Layoat Page 7
i l l 909190<A apex of the bend (caused by the manufacturing p'rocess) is carried over into the adjacent unstif fened straight tube. Additional conservatism is incorporated since this maximum ovality has been assumed to occur in conjunction with the helium maximum hot atreak temperature. j The creep subroutine of the SUCKLE computer code, Re f. 3, was modified to incorporate the creep equations for the five steam generator materials in the temperature range of this event (see Figs. 4.2 anc 4.3 for the creep subroutine). The tube dimensions, ovalities, material types, pressures, and range of temperatures investigated are shown in Table 4.1. The carbon steel inlet tube that connects the primary closure with the triturcated feedwater header was evaluated under this loading. The carbon steel tubes have thicker walls and, therefore, have a lower pressure stress. The maximum ovality of the carbon steel tube has been assumed to be 9%, the same as the economizer outlet tuce bends. Times to creep collapse were plotted as a function of tube temperature for each of the five materials [see Fig. 4.4(a) through 4.4(e)]. The tube metal temperatures were conservatively assumed to be the same as the helium hot streak temperatures. This analysis addresses a steam generator without cooling water in the reheater. The helium hot streak varies from inlet to outlet of the steam generator, due to mixing of the helium as the helium passes over the steam generator tube bundles. The creep buckling computer pecgram, BUCKLE, calculates the time for the tube ovality to attain the value which causes the tube maximum local stress to equal the material yield stress. The program uses constant values for temperature and external pressure. A parametric study using the parameters of Table 4.1 has shown in Fig. 4.4 that the most critical steam generator tube for creep buckling above 1330*F is at the economizer bottom end, i.e. , at the feedwater inlet. The economizer is manufactured from T2, a 1/2Cr-1/2Fb steel. At temperatures below 1330*F the carbon steel feedwater inlet tube becomes most critical. Therefore, both materials were evaluated. Page d
909190/A i FUNCTION ECREEP (CURTI , DT , TEMP , STRESS , STRAIN ,0S , IOP , II , MTYPE) C C DOUBLE PRECISION C IMPLICIT DOUBLE PRECISION (A-H,0-Z) C C CREEP EQ. EC=A=(STRESS ==AN)= (TIME ==GAWA) C MTYPE=1 A80001 C HICH TEMP C MTYPE=2 T2 C HICH TEMP C0 TD (100,200), MTYPE 100 U1=-1.3E4 U2=0. U3=1. U4=6.04E4 U5=-23.44 C0 TO 101 200 UI=-6.45E3 U2=0. U3=1. U4=4.164E4 U5=-19.9 101 T= TEMP C= (T-32. ) = 5. /9. TR=T+460. CAWA=1./(U2/TR+U3) X=1./CAMMA AN=-U1/TR=CAMMA A=10. = = (- (U4 102 IF (II.EQ.1)/TR+U5) WRITE (6,10) =CAMMA)T,C, AN, A,CAMMA 10 FORMAT (5X, 'C T A AN CAMMA',/,5X,5D16.8) B=(STRESS /1000.) AN 1 TBAR= (STRAIN =100./ (A.B)) = =X C0 TO (1,2), IOP 1 CONTINUE C CREEP STRAIN INCREMENT ECREEP=A B. ( (TBAR + 0T) = = CAMMA-TBAR = = C AMMA) /100. RETURN 2 CONTINUE C- TIME INCREMENT ECREEP= RETURN (DS = 100. / (A.B) +TBAR = = CAMMA) = = X-TBAR t 90 i Fig. 4.2. Creep Sabroutine for A800 Grade I and T2 Material
- at High Temperatures Page 9
l
. - og FUNCTION ECREEP (CURTI , DT , TEMP , STRESS , STR A IN , DS , I OP , II , MTYPE?
C C DOUBLE PRECISION l C 1 IIR.ICIT DOUBLE PRECISION (A-H,0-Z) C . C CREEP ER. EC=A.(STRESS..AN) (TIME.. CAMA) C WTYPE=1 A800C1 C HICH TEMP I C WTYPE=2 T2 C HICH TEMP C WTYPE=3 A800H C HICH TEMP C WTYPE=4 T22 C HIGH TEMP C WTYPE=5 SA 516 CR 55 C0 TD (100,200,300,400,500), MTYPE 100 U1=-1.304 U2=0. U3=1. U4=6.0404 U5=-23.44 C0 TO 101 200 U1=-6.45D3 U2:0. U3=1. U4=4.164D4 U5=-19.9 CD TO 101 l 300 U1=-1.21626D04 U2=1.07849D03 U3=0.47166 ! U4=5.141660004 U5=-19.0780 C0 TO 101 400 U1=-1.22004 U2=-1.169D03 U3:2.125 U4=5.64004 U5=-25.49 C0 TO 101 500 U1=-6.1468003 U2:0. U3=1. U4=2.794D4 U5=-12.5245 101 T= TEMP C=(T-32.) 5./9. TR=T.460. CAMMA=1./(U2/TR.U3) X=1./CAWA AN=-U1/TR.CAMMA l 102 A=10... (-(U4/TR+U5) IF (II.EQ.1) WRITE (6,10) .CAMMA)T,C, AN, A, CAMMA 10FORMATj5X,'C T A AN CAWA',/,5X,5D16.8) ! B=(STR SS/1000.)..AN l TBAR=(STRAIN.100. / (A.B)) ..X 6 CD TD (1,2), IOP 1 CONTINUE C CREEP STRAIN INCREMENT ECREEP=A.B. ((TBAR + DT) . . CAWA-TBAR. . CAMA) /100. RETURN l 2 CONTINUE l C TIME INCREMENT l ECREEP= (DS.100. / ( A.B) +TBAR. .CAWA) . .X-TBAR
- RETURN
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909'90/A ( TABLE 4.1 BUCKLE PARAMETERS FOR EES TUBES Temp. Wall Range Pressure Ovality 0.D. Thickne== Location Material *F psi % Inches Inchea Super-heater II Alloy 800 H 1,400 832 9&7 1.00 0.205 and 732 1,800 Dcwncocer Alloy 800 1,400 832 9&7 1.00 0.205 Grade 1 and 732 1,800 Evaporator T22 1,200 800 1.00 7 0.125 1,300 Econcmizer T2 1,200 800 9 1.00 0.125 (Feedwater inlet) 1,300 Feedwater Carbon steel 1,200 800 9 1.25 0.137 inlet tube 9 1,350 M
!!ote: Ovality - Ovality --
w ere D g - Mean diameter ASME 4 Page 11
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. . . i i ~9 8 8, 8, 8 8 3
, ~- ~ - $_
(3) e;npuedwaj. Fig. 4.4(c). T22 Evapor ator* Page 14
Time to Buckle for T2; 0.D. = 1.0 in 1400 & 9 a .
.+.
, , ? . 4 .
. +
, + g
- j '
gg. . . . . . . s s . t.
.9..-
.. .- ..---..;.--....5
..g..-...;...v.. ..t.-
, s .
4 : . 3 - *
- 4 33 .. ,..>.4.....g....4.. .- ...t.. ..~..p.~.- + + . .: -- 4. 4...:.. .g 4.- .+ .. ; . . t .
r .
! : . ? ? !, , .-
g . . . - ..
- m.
- 4 .
?
6 g g .. . . . _ . . .; .
- . p . .~
-..+.. . ..
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- ..+ .- 9 -q .. ..g.7. . . ; .. p . -
- 9. . 4 .. . . . t .. .
.C A
; * ? . . ,
j j k . 6 f . . c , .
...e V . . . . .
^ .+...,. - .t.-..,. .. . . - . &.- 4. :- ; .4..+4......,;.....+. .
. . . . . . t.. ;
a e .-
-. u : . , . .-
5 . : g -1 . , 3_ , - : : . . . . . : 4 o gg_ .
- y. .~. . . . _. _ ; .. 9. -: ..
9 9..... 2..e.
, ;. p... . r . . .. ..
. 9 .. ., .. ... y .. r .. . ..-7.-....
a 6. . 4 .t : . m e . : + . O. . : . . m c . . . .n._.... .. . . .
....c ..+- r _ .~ - . - - . . . .. e. . .. . ..r a
4.+.... . . . , ..s pe g . g
~
O 8 . . ' * . . . *
- p.= t . 6 .
3
. . : 4 : .. .
. . . 4 . . . :
. , 6 -, ,
r- r- * -- -'---*----*'------*-:--:**-+--::2v-'- -+- O GOO - -----r-----r- --- * +-+--t-m* j 3 : . : . * - ,
*
- g r
N : - t
- l, . .? ,
? :
.. . - . . .p...,.
G .._.....;--
".. t 7.+.. ,..t.. - - . . ..p.+.-.-,-..:.....:- 4...t. . t .. p ..J .- . .-. ..t.
*'3 . : . -
.e ' i .
* ? ' 4 ,
; I :
- gg_ ..-....;..--.. ..
...).-4..,
.y-;. 4 ..
....e.,
. . - -~ - , + . .
. . . . . t - r. +: . .t.-
..).-<.
t *
.
- I : 4 .t 4
! : . e :
..s..+....r....,..
- . + . . . . . . . ..&.
. . . . . , .. . . . . . , _ . . .7
. . . . . u. . . ,
- . : 4 . , ;
- 190 0 . . . . . . . . . . . . . . . .
10' d 10' A ute (hr) P = 800 Psi; Ov=97. {e"
" Q
- N lL*
1 l l Time io Buckle for Corbon Steel; O.D. = 1.0 in woo n
- m. - ~- , .. .
e i c 1350-l i
^
w
* \ - .
I o, c
~1 1300- . - > .- --
E ^ a b - - u, , . e '
@ 3
\ r n o nu- . . .
.y e C E o E -~ -
a
- e os o 1200- - -
g - ~ - i 115 0 - - - - - ! a ! u or G
'n a s a s , , , , , , , ,
y 5.0= Xf d d g g Ufe (hr) $ e . m P = 800 Psi; Ovr9"/. % 3+
909190 / A When the initial overly conservative estimate
- of tne maximum helium temperature and economiaer tube pressures were incorporated into the BUCKLE code, the time to buckle was le=s than the event time.
However, the estimates of the temperature and pre 992re showed that they varied significantly with time. The method employed to conservatively assess the likelihood of creep instability involves using the BUCKLE code constant-pres *ure results to obtain creep buckle times for di*fering pressure di "eren-tiala. To account 'or a varying external pressure histcry, the premmare was divided into convenient constant pressure levels P n
' "Ud Dh' DI"' "D this pressure level was t. The time to creep collapse T n wa* found n
using BUCKLE for each pressure level at its corresponding temperature. Using the aummation o' t /T n n wa* not correct, because all the Tn value* were based on the *ame initial ovality. The value= of ovality were not available ' rom the BUCKLE code, ao two assumptions were made:
- 1. For a given constant temperature and pressure, the critical tube ovality where the tabe strea9 im equal to the material yield stress is independent of time.
- 2. For a given constant temperatare and pressure, the rate o' change o' ovality dJe to creep varie* exponentially with time.
For the lire aammation method de*cribed below, this la a conservative aasumption. Theae aamumptions allow the ovality changes with time to ecnaervatively be predicted under conditions o' varying pres *are and temperature. Failure la predicted to occar when the ovality at a specific temperature and pre **;re reaches that value which caume9 the atreas in the tube to reacn yield atreas. Page 17 i . . . . .
90 9190 ' A As ahown in Fig. 4.5, linea repreaenting evality changea with time are drawn #or each presaure/ temperature combination. Theae lines were drawn from the initial ovality at time zero to a point defined by the ovality to cause instantaneoua backling and the time to reach creep buckling. For a specific pressare/ temperature hiatcry, ovality changea are accounted for by dividing the load history into diacrete time periods t of n constant pressure /te=perature combination. The t tim' n period values were superimpoaed in the correct aeq...nce on the creep rate lines. The 'inal ovality at the end of time period t w? he n starting ovality for time period O +1 n Fig 2re 4.5 waa uaed to per'ccm the aummation o' nt which took into account the e#fective creep time lost between starting ovality at the beginning o# time period tn and the ovality at end o' time period t n-1* In crder to determine the maximum temperature of helium that the feedwater inlet tube to the econcmizer can withstand during this event, the following simpl1#ying assumptions were made:
- 1. The helium presaure-time hiatcry supplied on November 5, 1986 (aee Fig. 4.6) doe
- not change with amall variation = o' helium temperature.
- 2. The shape o' the temperature-time hiatory mapplied on November 5, 1986 (see Fig. 4.7) does not change when the amplitade la varied.
3 The maximum temperature o r the T2 material will not exceed 1350*F. U*ing theae asaamptiona, a parametric at 2dy (Table 4.2 ahowa the variable *) c' the e #ect r o* varying temperat2re for the pre =* area c e Fig. 4.6 wa= perfccmed. Table 4 3 showa which temperat 2re and preaaare ccmbinaticna were evaluated. Thia table alao showa the calculated craep b20kling timea 'or the conaldered prema2re-temperat2re combinationa. Page 18
909190/A I I l l l Positive i
, Margin 1
p' f3 Failure ovality
$1 /
1 ?
/
/
ds
/ L
/
/
w l/
$ s
% /
5 /
/ , . ' P,
# Initial Ovality
=
6 = 4 i Time to Failure i Fig. 4.5. Backling ovality Vertaa Time to craep Collapaa Page 19
I 909190 / A l l l l l l ~ ho l 1 e
}
i 1 I
- m Ia m w
! = . I. 3 I
- I
'I M
A E 3 ar
's t N
! [
t
*:, ****, e n s ,, . ,,
.,,, e I - :
j Am-c Fig. 4.6. Helium Pre *a re History Page 20
SG fl0D0LE ItstET TENERATUEES - 35 ru Flou uro LiteER C00Lisec 1400 LEGEND HOT 2 T E 1300
~
/.
7
\ -
" "A6'd - "
? N
, p -
L
' E
~
g - g n 1200 -
; T li U -
G '. a n R E l i 1100 : g - '.-
% D - - ~~. .
E - G - ~~~. . Q ' S Q F 1000
%,W..
900 i i i O 2 4 6 8 10 12 8 4.. TIME, 1100RS G a ~. y 3'*
90 9190 /A TABLE 4.2 PROPERTIES USED TO CALCULATE THE CREEP BUCKLING COLLAPSE TIMES FOR THE EVAPORATOR TUBE Temperature E psi x 10' 3) a_ 10
- Poissog Yield {gess
'F Ratio ksi 1.350 8.4 0.808 0.236 7.5 1.300 10.8 0.804 0.243 10.0 1,250 13.2 0.800 0.250 12.25 1,200 15.6 0.797 0.257 14.5 1,150 18.04 0.793 0.264 16.25 1,100 20.48 0.79 0. 271 17.5
" From Ref. 13 From Ret. 14 Page 22
909190/A TABLE 4.3 TIME TO CREEP BUCKLING COLLAPSE (h) Temperature, *F Pressure Psi 1350 1300 1250 1200 1100 832 5.00 800 6.3 23 7 87 3 760 7.92 28.0 103 750 8.33 29 3 41 4 740 8.75 30.6 III 730 32.0 115 448 720 120 467 9930 710 487 >10000 700 507 >10000 l l l '. Page 23
. \
909190/A The ovality to cause the tube surface to reach yield was calculated using a linear relationship between ovality and o /P, wnere o is the y y yield stress at temperature and P is the applied pressure. Two catum points were determined by the BUCKLE code. Both Refs. 7 and 8 show that the relationship of ovality to o /P is very close to linear. These y , values are tabulated in Table 4.4. Using curves of the collapse ovality and the time for creep collapse in Fig. 4.8, the economizer tubes were determined not to suffer creep collapse failure during the postulated firewater cooldown event 3 from 35% power extrapolated to a 1350*F maximum temperature. The
; pressure event history from Fig. 4.6, together with the temperature a
history from Fig. 4.7, were idealized into the following loadings: 4
- 1. Pressure of 832 psi e 1350*F for 1 h.
- 2. Pressure of 760 psi @ 1300*F for an additional I h.
3 Pressure of 760 psi @ 1250*F for a further 1 h period. t Using Fig. 4.8 to account for the starting ovality of the second and third loadings, the predicted collapse times for these three idealized loadings were determined as:
- 1. Collapse time of 5 h ending ovality after 1 h is 9.0%.
- 2. Collapse time of 26 h ending ovality after 1 h is 9.0%.
t 3 collapse time of 102 h, ending ovality after 1 h is 9.1%. This leaves an equivalent time remaining to cause creep buckling of
>400 n, assuming the pressure remains at 750 psi and the temperature
- remains at 1200*F.
l For the carbon steel tubes the collapse times were calculated using the BUCKLE code, Ref. 3, for the three idealized loadings: i i i l Page 24
909190 /R I TABLE 4.4 i TUBE OVALITY TO CAUSE INSTANTANEOUS BUCKLINO COLLAPSE (%) Temperature, 'F Pressure psi 1350 1300 1250 1200 1100 832 9.01 15.7 21.7 800 9.8 16.8 23.0 760 10.9 18.2 24.8 750 11.2 18.6 25 3 31.9 740 11.5 19.0(*} 25.8 32.5 730 26 3 33 1 42 3 720 26.8 33.7 43 0 710 34.4 43.8 700 I 35 ") 44.5
- Determined using the BUCKLE code, a
Ovality % = b -l - 4.96 45 P l l Page 25 l
Time to Buckle for T2; 0.D. = 1.0 in 40
~ . . , . . ... .. .. . . . . . . , . . . .. . . .
u
- c. .
o
#9 K . . . . . , ..q... .. g. . ..,....y.. ,
..p.. , ; ,
- m. .3
'N
.s: c .. . . , . .
. ...;.. .. .. .'l m < -
- u -
4 >- v
. c, !.
x ^
-4 <
g . v
>. Legend
!! 4 .
u o .v.
- n c -O
, a A P=750 Psl m >
c' () O . . . . . . , .. . . . , . . . O P=730 Psi m . . . , . . . . . ..,.. ..,.. . . , . . . . 3 g . - .. . . o .:. O P=720 Psi o
,., O P=710 Psi ro .o. ! V P=700 Psi o e-a w O T,}
n v. Lt 10 -
?.
,y f.. .. . . .s' ..p.. ........i....
. , . .. ..p.
3.... ....g. ...... ..,....p....,.. . . . .,..,,..p. i . 8 i i i . . . . . . . 3.0 50.0 100.0 15 0.0 200.0 250.0 300.0 350.0 400.0 450.0 500.0 550.0 o so a W Life h c) e-(D r, T = 1200 I sg f 'S
- a. ~
g e k-u i 0 1 D. O 2
=
n p
. +
L is l d P Pa Ps Ps Ps n0 g = = 7= = = e P P P P P AO O.O V i i e 0807 604 30 2707 l
,0
. 1
. 0 0
0 2
. 10 0
0
. 8 r
D 0 5 2 1 I
=
g e i T .
. 3 .
r f o
)r l
e . ' . . 0. (b k i0 e c 6f i u L B 4 o
, ; l-t e .
. 0 i
m 4 i0 l 4
. 0 0
2 7 : 0
.0 3 8 0 mMvI._S P hS. dj 3, % N D 6 e .
. 9m y g $
l'
i 909190/A i l l l l 6 O G) C O O m o - d il m , mE2222 58ggo o ; o ED N n R% -
.O c ,
e LLLLL a
;.. <e o ac>- :
o
; -Y 1I --
4.-
.n 3
o - d.
+- .
cN b _o u g + o -n T n
@ C 2O 4 e
- =
a a co . .o
.e o -
i - . .E w - o i t _
.g
.! - .. .o
.o o
2 e = (%) 4!P^0 Fig. 4.8(c). Temperature - 1300a? Page 28
909190 / A w i a ! O m ]
.c _ _ -e*
o d u222ff 5o 8884 h o 6 vao.<
~
O ht.
-3 & T .9 i.
3 ob f *$" 3 CD M - o E F -y
-b
-s o.
N o m o (*4) A i!! D^ O l l Fig. 4.8(d). Temperat 2re - 1350of Page 29
Time to Buckle for T2; 0.D. = 1.0 in 820 , i
- 4 i i
- e
~
< , , <- - l.
Legend y . .t .. ..; .
? 800- ..,.
~*- "T* .. .. ., .
"I. T
.! V T = 1350 Dea F
,e ; t
. . . .. ; .. .. ;. .. : . . -i- i O T = 1300 F T
.3 "1" " ! S T = 1250 Dea F
- 780- "i" '
"i" - * " '
"i" I
- i
- }
"O T = 1200 Dec F
. . f. . .. t. . . i. . .. i.. +.. 1' r I- .
-~! - --
-t rti t -
~+"
,m n , !
m M
- a i- -}- I ?-
vOL 760- + , , i . 4. .
"g" 4 :
I . i 2 4 4 . . . . . . ..+.! I, i.. v- .>- i- i- -"9- -i- ~!' e 3 -
. t s
M , . i n g_ ..g..,- ..<-
-.s.. . . ,.. .i.. .
ct u e a o I n ..a.. ...r.+. 4 y . . ;. .
. . i.. .j . i, . ,
- < 4- >
+-!-r- !-
O ' H
+
i m . . . a .
. .; .. '.1.. .
- j. .: .. . . .
720 - '"!."*""+" .
'~'"",'
rt : r ; * . o .--a-p .. t. . . . ,
..i. 4.1. .
- i. .
,-t- -
o .
+ +
+
y
! 4 f~ "i" 700- ~ 5" * ' " 1" 5" "I'+ "^"
"O "?" " I"
, . . ;. . . . ..p i+ .j.+ + t-
'i -
' - ~-
; ~!"
. t t
680 ' ' ' '
d i a ' +1 iaa a e ' ' '
u if id 'd i 1 y Life (hr) o
<D $
w O Ovality = 9% g s
909190/A
- 1. Collapse time of 7.6 h for 832 psi pressure at 1350*F.
- 2. Collapse time of 20.1 h for 760 psi pressure at 1300*F.
3 Collapse time of 45.4 h for 760 psi pressure at 1250*F. These values show that the carbon steel feedwater inlet tubes will not collapse for the considered loading. A further considerable conservatism exists since, even if the tube does undergo creep collapse, the pressure boundary is not likely to be ruptured, because the residual creep ductility of the material is substantial at such high temperatures. The tube could probably collapse to a flat strip and not rupture the material.
- 5. REHEATER HELICAL BUNDLE A structural evaluation of the reheater helical bundle under thermal expansion loading was deemed prudent due to the possibly large elastic follow-up action of tubing adjacent to a region where local yielding has occurred. The local yielding will relieve the overall stress levels, but the elastic spring action of the adjacent portions of the helical tube could cause surface strains in local areas to be excessive. The major load contributor to the tube stresses is caused by the relative radial displacement of the tube and support structure. A portion of the tube support structure will be at the hot helium tempera-ture, while the tubes will be relatively cold due to the 80*F firewater inside of them.
The analysis of the reheater helical tube bundle used the MODSAP (Ref. 9) finite element computer code to evaluate the tube forces. The reheater helical tube model used was one developed previously. Due to the higher spring rate of the smaller diameter coil, the innermost helix is tne most critical from the effects of support plate displacements. From a previous unpublished evaluation, tube No. 82 was found to have Page 31
4 909190/A the highest bending stress under normal steady-state thermal loading = at 25% and 1001 power. The temperatures used for this analysis were taken from Ref. 1,and Appendix A of this document. Tube support structure temperatures, other than those in Appendix A, were assumed as either at the water tempera-ture or the helium temperature, depending on the location; i.e., the headers were assumed to be at water temperature as they are insulated
; from the helium, and the support structure slotted region was aasumed to i
be at helium temperature. The values for material properties were taken from the ASME code, Ref. 2. The initial elastic MODSAP computer model was expected to show regions where the reheater tube stresees exceed the yield strength of the material. The hottest of these high-stress regions was then modeled as having a plastic hinge. That is, a hinge was incorporated into the model and a moment applied across that hinge equal to the maximum mcment that the tube section could produce during plastic deformation, aasaming no kinematic hardening of the material. The equations used to calculate the fully plaatic moment were taken from Ref. 10. An additional . plastic hinge was later incorporated in the model as the yield streas was still exceeded in other parts of the model. The effective length of each plastic hinge was taken as the thickness of the perforated sspport plate, because the tube is locally hotter there due to the fin e*fect of the plate. The total bending strain at the plastic hinge was calculated uaing the hinge rotation from the MODSAP outpat and applying that rota-tion geometrically to a 0.75-in. length of reheater tubing. Thia , 0.75 in. corresponds to the reheater support plate thickness. Using the average strain over the 0.75 in. length is an acceptable procedure aa atrain hardening will not allow the strain to occur over a very emall length. Additional computer runs were made to determine the effects on the reheater tubes o' the collapae of the tube support structure. Thia w a *. Pace 32
909190 /A deemed necessary, since part of the support structure will be at helium temperature and could experience a high compresaive load which might cause a creep buckling "ailure mode of the support structure. Only one reheater tube was evaluated. This was tube No. 82, which is located in helix No.1, the innermo9t helix. This tube is modeled aa a string of finite-length, curved pipe elements; the geometry coordi-nates are also shown in Fig. 5.1. This model was input to the MODS AP (Ref. 9) computer code to evaluate the local loads on each pipe element under the applied loadings. The applied loading consists of a tabe l temperature and a support structure displacement are also included in Fig. 5.1. The displacement of the support structure is cauaed by the thermal expansion of the headers, supports, and tne perforated aupport plates. The datum for the vertical displacement calculation la the bottom edge of the support plate, but the MODSAP vertical datum is the center line of the tube at the lower nozzle (see Fig. 5.2). There are nine locations where the tube is subjected to structural displacements (boundary conditions "or MODSAP). Two of these are the input and output tube nozzles, and the other seven are where the helical tube passea through the three perforated support plates (see Fig. 5. 3) . The temperatures of the tube supporting components are dependent en both helium and cooling water. The headers are insulated and there#cre were amaumed to be totally dependent on the cooling water temperature. The displacements of the headers are shown in Table 5.1. The perforated tube aupport plates are heated by the helium and cooled by cloae contact with the reheater tubes (9ee Appendix A) . The temperature of theae plates is shown in Table 5.2 as a function co vertical poaition. The temperature variation is cauaed by heat removal from the helium by the cooling water in the tubes. The temperature of the atracture between the headers and the perforated plate 9 will be af#ected by the tempera-tures o* botn theme componenta and the temperature of the heliam. The Page 33 t
CASG,A F#CDTUBE820= PLOT. ';- a .y ^i . QUSE 14.,FWCDTUBE820= PLOT. Oprt fsv-reheat.f.cdtube820 cadd smb. prog.modsap/ execute 52 2 FSV1 REHEATER TUBE HELIX 1 FIREWATER C00LDOWN (TUBE i 82) 1 1 1 1 8.173 .000 4.719 80. 2 12.990 .000 3 7.500 80. 17.164 3.819 4.654 120. 4 13.838 7.509 -2.319 127. 5 13.339 6 7.826 -3.365 128. ] 12.086 8.056 -5.765 129. 7 -6.695 10.000 -11.597 137. 8 -13.391 11.235 .000 142. 9 -6.695 12.469 11.597 147 10 6.695 13.703 11.597 153. 11 13.391 14.938 .000 159. 12 6.695 16.172 -11.597 165. 13 -6.695 14 17.406 -1;.597 173.
-13.391 18.641 .000 180.
15 -6.695 19.875 11.597 189. 16 6.695 17 21.110 11.597 7 99. 13.391 22.344 .000 209. 18 6.695 19 23.578 -11.597 220.
-6.695 24.813 -11.597 228.
20 -11.500 21 26.681 6.860 228.
-9.095 27.065 10.464 228.
22 -8.470 23 27.369 11.438 225. 24
-4.907 30.383 16.821 185.
.400 34.165 14.563 117.
25 1 1 .400 26 1 1 34.165 9.001 117. 1 1 9.173 27 1 1 1 1 0.0 4.719 1 1 1 1 8.173 28 1 1 0.0 5.719 1 1 1 8.173 29 1 1 1 1.0 4.719 1 1 1 -6.695 30 1 1 1 10.0 -12.595 1 1 1 -7.695 31 1 1 1 10.0 -11.597 1 1 1 1 1 -6.695 11.0 -11.597 32 1 1 1 1 1 -7.695 33 1 1 12.469 11.597 1 1 1 1 -6.695 34 1 1 12.469 12.597 1 35 1 1 1 1 -6.695 13.469 11.597 1 1 1 1 1 1 14.391 14.938 0.0 36 1 1 1 37 1 1 1 13.391 14.938 1.0 1 1 1 1 38 1 1 13.391 15.938 0.0 1 1 39 1 1 1 1 -7.695 17.407 -11.597 1 1 40 1 1 1 1 -6.695 17.407 -12.597
.1 1 1 1 1 -6.695 18.407 41 ~1' 1
-11.597 42 i 1 1 1 1 -7.695 19.875 11.597 1 1 43 1 1 1 1 -6.695 19.875 12.597 1 1 44 1 1 1 1 -6.695 20.875 11.597 1 1 1 1 45 1 1 14.391 22.344 0.0
'1 1 1 46 1 1 1 13.391 22.344 1.0 1 1 1 47 1 1 1 13.391 23.344 0.0 1 1 48 1 1 1 1 -6.695 24.813 -12.597 1 1 49 1 1 1 1 -7.695 24.813 -11.597 1 1 E0 1 1 1 1 -6.695 25.813 -11.597 1
El 1 1 1 1 1 1 -1.3999 34.165 9.0007 1 1 1 1 1 .3999 34.165 10.0007 E2 1 1 1 1 1 1 .3999 35.165 9.0007
=.g. . n :e v. - ~ :s voi,L :<; n : L :n Page 34
909190 /A 7 27
- 0. 1. O. O.
I 26 1 0.0006 0.0 1.0 O 1 27 1 0.0004 0.0 1.0e+08 1 28 1 -0.0005 0.0 1.0e+08 7 29 1 0.0692 0.0 1.0e+08 7 30 1 0.0399 0.0 1.0e+08 7 31 1 0.0049 0.0 1.0e+08 9 32 1 0.0483 0.0 1.0e+08 9 33 1 0.0837 0.0 1.0e+08 9 34 1 0.0090 0.0 1.0e+08 11 35 1 0.1134 0.0 1.0e+08 11 36 1 0.0 0.0 1.0e+00 11 37 1 0.0137 0.0 1.0e+08 13 38 1 0.0650 0.0 1.0e+08 13 39 1 0.1127 0.0 1.0e+08 13 40 1 0.0191 0.0 1.0e+08 15 41 1 0.0734 0.0 1.0e+08 15 42 1 0.1271 0.0 1.0e+08 15 43 1 0.0252 0.0 1.0e+08 17 44 1 0.1635 0.0 1.0e+0S 17 45 1 0.0 0.0 1.0e+00 17 46 1 0.0323 0.0 1.0e+08 19 47 1 0.1561 0.0 1.0e+08 19 48 1 0.0901 0.0 1.0e+08 19 49 1 0.0405 0.0 1.0e+0S 25 50 1 0.0001 0.0 1.0e+08 75 51 1 0.0034 0.0 1.0e+08
'5 52 1 0.0055 0.0 1.0e+08 2 24 1 8 1 1 8 INCOLY 800 FROM CASE 1592 70.0 28500000. 0.339 0.00000775 100.0 28400000. 0.339 0.00000795 200.0 27800000. 0.341 0.00000834 300.0 27400000. 0.347 0.00000860 400.0 27100000. 0.353 0.00000878 500.0 26600000. 0.355 0.00000892 600.0 26400000. 0.357 0.00000900 1250.0 22100000. 0.379 0.00000965 Fig. 5.1. (Continued)
Page 35
909190/ A 1 1.125 0.140 0.13 0.0 0.0 0.0 0.0
-1, 0.0 0.0 0.0 0.0 0.0 0.0 0.0 0.0 1.0 0.0 0.0 0.0 0.0 0.0 0.0 1T 1 2 1 1 70.00 100.00 2B 2 3 1 1 70.00 100.00 3.375 CC 14.283 2.169 5.261 4000 3T 3 4 1 1 70.00 100.00 4B 4 5 1 1 70.00 100.00 3.375 CC 13.212 4.464 -3.632 4000 ST 5 6 1 1 70.00 100.00 6B 6 7 1 1 70.00 100.00 13.426 CC 0.000 9.028 0.000 4000 78 7 8 1 1 70.00 100.00 13.405 CC 0.000 10.617 0.000 4000 88 8 9 1 1 70.00 100.00 13.405 CC 0.000 11.852 0.000 4000 98 9 10 1 1 70.00 100.00 13.405 CC 0.000 13.086 0.000 4000 10B 10 11 1 1 70.00 100.00 13.405 CC 0.000 14.321 0.000 4000 llB 11 12 1 1 70.00 100.00 13.405 CC 0.000 15.555 0.000 4000 123 12 13 1 1 70.00 100.00
.:.405 CC 0.000 16.789 0.000 .4000 13E 13 14 1 1 70.00 100.00
- 3.405 CC 0.000 18.024 0.000 .4000 14E 14 15 1 1 70.00 100.00 13.405 CC 0.000 19.258 0.00 .4000 158 15 16 1 1 70.00 100.00 13.405 CC 0.000 20.492 0.00 .4000 168 16 17 1 1 70.00 100.00 13.405 CC O.000 21.727 0.00 .4000 17B 17 18 1 1 70.00 '^0.00 13.405 CC 0.000 22.961 0.00 .4000 188 18 19 1 1 70.00 100.00 13.405 CC 0.000 24.196 0.00 .4000 198 19 20 1 1 70.00 100.00 13.423 CC 0.000 25.747 0.00 .4000 20T 20 21 1 1 70.00 100.00 21B 21 22 1 1 70.00 100.00 3.375 CC -9.258 30.427 10.247 .4000 22T 22 23 1 1 70.00 100.00 23B 23 24 1 1 70.00 100.00 3.375 CC -2.985 31.996 14.563 4000 24T 24 25 1 1 70.00 100.00 l
0.0 1.0 0.0 0.0 Fig. 5.1. (Continued) Page 36
909190/ A l l I
~
f %
% '. J l
l Reheater Upper Collector I Helix 1l Tubes T- Aqg
# I
#r ! %*
# l
#- . L___-
Plate "Y" REF $ 1 ie- 7T i (55.S) k
*10DSAP "Y" REF H911x 1 Tubis ."/S[/f,2g)-> Y%
- 7. .
I j Fig. 5.2. Reheater Support Plate Dimension = l Page 37 I
909190/A l r ~~ < L . - y. 1 p Upper Collector p f Perforated
?
8 -'
,, Support Plate
\ *
- I
\
f h t fm
.15 3__ E :
f lj .. ' ,r s. . %}, l7 "' %
} -}
[ \
\
'. /7 9 ,
3( i,/ M" , * .
~
s' y N u . ( ' \ i V:~ , 7 i . . -
. O Slotted Support P
'l
- Q srun /. '
- HEL / 3'M NODE I i DE ,
I j~-
/' ser .:nt,, , . .
Fig. 5.3 Reheater Helical Tube Model Page 38
f 90 9190 /A TABLE 5.1 (1) RADIAL THERMAL DISPLACEMENT OF HEADERS AT HELIX 1 N0ZZLES 6 Radius a v Ncde Location In. (R) T 'F In./In. 'F Ra(T-70) No. Water 9.5 80 7.82 x 10
- o, coo 7 i Inlet Y - 0.00
~
Water 9.1 117 8.02 x 10 ' O.0034 25 Outlet Y - 34.164 (2) VERTICAL THERMAL DISPLACEMENT OF HEADERS AT HELIX 1 N0ZZLES Location Y In. T 'F aY T F
- Ave In. In. 'F 6
7 [o y Node No. 1 Water Inlet 0.0 80 Bottom of 6.75 80
~
7.82x10 ' O.0005 -0.0005 Plate 6.75 80
~
1.42 80 7.82x10 ' O.0001 0.0004 8.165 80
~
8.0 86 7.86x10 ' O.0010 0.0006 16.165 91 8.0 97 7.93x10 O.0017 0.0023 24.165 103 - 10.0 110 7.99x10
- 0.0032 0.0055 Water 34.165 117 Outlet Node No. 25
- From bottom of plate.
Page 39
909190/A TABLE 5.2 REHEATER SUPPORT PLATE TEMFERATURE Distance from Top Temp. MODSAP (Inches) *F "Y" Location 3 0.0 473 25.07 1.176 449 23.89 2 352 425 22.72 3.528 403 21.54 4.704 383 20.37 5.88 364 19.19 7.056 346 1 8.01 8.232 330 16.84 9.408 31 4 15.66 10.584 299 14.49 11.76 285 1 3. 31 12.936 273 12.13 14.112 261 10.96 15.288 249 9.78 16.464 239 8.61 17.64 229 7 43 Helium Temperatures Ave. Inlet = 1307aF at top Ave. Outlet - 458aF at bottom Water Temperatures Inlet = 80*F at bottom Outlet - 117'F at top Page 40
909190 / A equations used for the slotted support structure radial displacement
- are:
Let a = B-C b = 0-B ee ig. 5. and Table 5. 3(2) . c = 0-C a = 60' Then, using Cosine rule: a2 = b 2 +c 2
- 2bc Cos a Since Cos 60 - 0.5:
) 0 = c2 - bc + (b2 - a*) , a qJadratic equation 2 a B (B - 4AC) 2A For 0-B = 10.0 in. 0-C = 11.257 in. 0 (100 - 4(100 - 11.2572 )]*/ 2 c = 2 e c = 12.192 in. @ room temperature For 6L(B-C)
=
0.1134 in. c = 12 373 in. @ top 6L(B-C) 0.0330 in. e = 12.248 in. @ bottom Page 41
909190/ A Perforated Suppcre Plate p%/ Helix 1 Tubes Y C Slotted Support Plate e I e x \
/
/
/
\ .,/
0 (, J
/
i Fig. 5.4 Plan Diagram o' Reheater Sapport Stract are Page 42
90 9190 / A TABLE 5 3 DISPLACEMENT CF SUPPORT STRUCTURE ( 1) '/ERTICAL Y Temp. 6 aY T a v OA + [6 v MODSAP In. 'F In. In./In.*F ayo(T-70) Node No. (a) (b)
~
OA(c) 112( } 2.25 80 7.82x10 ' O.0002 0.0002 l I 6.75 229 ~ 3.25 240 8.44x10 ' O.0047 0.0049 7 l 10.00 251 ~ 2.47 263 8.50x10 ' O.0041 0.0090 9 12.47 275 2.47 290 8.57x10 O.0047 0.0137 11 14.94 304 _ 2.47 321 8.64x10 ' O.0054 0.0191 13 17.41 337 _ 2.47 356 8.64x10 ' O.0061 0.0252 15 19.88 374 _ 2.46 398 8.78x10 ' O.0071 0.0323 17 22 34 421 2.47 444 8.84x10
- 0.0082 0.0405 19 24.81 467 i
Measured from bottom of plate. (D From Table 4.1. U See Fig. 4 3 (d)Within header insulation. Page 43 l ___ _ _ _ _ _ _ _ _ _ _ _ _ _ ._ _ _ .
909190/A TABLE 5 3 (Continued) (2) RADIAL TO HELIX 1 Component Length T gy, a 6 g 6 R
~
Top 0-A 8.0 117 8.02x10 ' O.0030 0.0015
~
A-B 2.0 117 8.02x10 ' O.0008 0.0004'
~
B-C 11.26 1,127 9.53x10 ' O.1134 0.1805
~
C-Tabe 1.2 467 8.87x10 ' O.0042 0.0042 Y = 25.75 {0.1866
~
Bottom 0-A 8.5 80 7.82x10 ' O.0007 0.0004
~
A-B 1.5 80 7.82x10 ' O.0001 0.0000
~
B-C 11.26 404 8.78x10 ' O.0330 0.0559
~
C-Tabe 1.2 229 8.42x10 ' O.0016 0.0016 Y = 6.75 10.0579 6 ~ * * ~ U8 RO-A L RAB L
+
6 R-B-C
= - -
(0-B)2-((B-C)+ Sg P} - 12.192 2 2 , 6 R B-C or 6g =0.1134)=0.1805 6 R B-C r6 g=0.0330)-0.0559 Fce input displacements to the MODSAP code, uae linear interpols-tion between Y = 6.75 and Y = 25.75 in. Page 44
909'90/A Then 6 - 12.373 - 12.192 - 0.181 in. @ top R(B-C) 6 - 12.248 - 12.192 - 0.033 in. @ bottom R(B-C) Linear interpolation was used to obtain the radial diaplacements o' the tube nodal displacementa at '.ne junction with the support plates. The thermal displacements of the support structure due to "ree thermal expansion are calculated by hand in tabular form in Table 5.1 for the header displacements, and Table 5.3 for the support structure displacements. The combined displacements are summarized in Table 5.4 The tube temperatures are in Appendix A and are aleo shown in Table 5.5. The reheater tube stresses are expected to exceed the yield streas limit of Saniero 31. This yielding will be local and will occur at a support plate. To evaluate the local aurface strains that the tube will be subjected to, hinges are incorporated into the model at the high Stress locations. A moment loading is placed across these hingea to represent the maximum moment the tube could develop a*ter exceeding the yield stress, but without any kinematic hardening o* the material. Thia maximum tube moment, the plastic moment, is calculated using the equation of Ref. 10. Mc M The tube 9 tress c - I Z To evaluate the limit moment the a ia made equal to the yield atreaa o# the material at the mean temperature o r the tabe. Page 45
909190/ A TABLE 5.4 MODSAP DISPLACEMENT
SUMMARY
HELIX ' TUBE 82 MODSAP Angular Displacement. In. Node Position No. f G Dcg. X Z R=(X8 +Z 2
)i a y 1 30 0.0006 0.0004 0.0007 -0.0005 7 240 -0.0399 -0.0692 0.0799 0.0049 9 120 -0.0483 0.0837 0.0966 0.0090 11 0 0.1134 0.0 0.1134 0.0137 13 240 -0.0650 -0.1127 0.1301 0.01 91 15 120 -0.0734 0.1271 0.1468 0.0252 17 0 0.1635 0.0 0.1635 0.0323 19 240 -0.0901 -0.1561 0.1803 0.0405 25 92.5 -0.0001 0.0034 0.0034 0.0055 X =
R Cos 0 Z = R Sin 0
~
Page 46
l 909190 /A l TABLE 5.5 REHEATER TUBE TEMPERATURE AT 50% PCWER Distance From Top Tube Temp. AT MODSAP (Inches) 'F 'F Model "Y" Water out 117 0 34.17 0.00 228 128 27.07 0.588 228 128 24.34 1.764 217 118 23.17 2.94 207 110 21.99 4.12 197 102 20.81 5.29 188 95 19.64 6.47 180 88 18.46 7.65 173 83 17.28 l 8.82 165 77 16.11 ! 10.0 159 72 14.34 11.17 153 67 13.76 12.35 147 62 12.58 13.52 142 58 11.41 14.7 137 54 10.23 15.9 132 51 9.03 17.06 128 48 7.87 l Water In 80 0 0.00 Note: Uae linear interpolation for intermediate valuea. Page 47
909190/A i Z, the section modulus, is given by: i 1. i Z = (D'-d') 32D 1 and the plastic section modulus Zp , is given by: D: [ Z = -- 1- { 1 - 2t- , (Ref. 10, Table 1.1) P 6 \ D 1 For the reheater tube, the specific values are: D - 1.125 in. I d - 0.845 in. t = 0.140 in. ) Z - 0.0953 in.' Z p
=
0.1367 in.' , At 200*F mean tube temperature a - 22,500 psi . 7 1 The plastic moment (i.e., the maximum moment the tube can resiat) ia: 1 1 M P 3075 in. lb . 1 1, ! This plaatic moment is applied as a constant value across the hinge in the model, to represent the maximum moment a tube can develop. 4 The ela9 tic tube model without any plastic hinges showed tube 4 bending moments exceeding the value of -the plastic moment M p in many elements. The higheat tube bending moment occurred at Node No.17, which_ is located towards the upper end of the tube bandle. A hinge with i moment PM waa incorporated into the computer model at Node No. 17 and t
) Page 48 J
t
- - - , - , , , , - , . - - - - - - - , , . . , - , ,---.no, -y .,, . - ~ , ,.--r-~ ,,. , - - --,,,- ,-..,--,,,-,a-, -
n--
4 909190 /A the bending moments and hinge rotations recalculated. The hinge rota-tion Oywaa 0.1855E-2 radiana. The tube surface strain was calculated using a tube outside diameter D of 1.125 in, and an effective hinge length L, of 0.75 in, c - Y (1005) = 0.14% 2L, Although the calculated strain is an average value over a 0.75 in., the strain hardening behavior of the tube material ensures that the strain will not become any more localized. If the radial force between the tube and plate could restrain the local compression yielding on the contact side of the tube, then the I maximum Strain on the other side of the (tube under tensile stress) is doubled to 0.28%. With a hinge at Node No. 17, the tube bending moment at Node No. 11 exceeded the plastic moment. An additional hinge with moment M waa p incorporated at Node No.11 and the bending moments and hinge rotations { again recalculated. The hinge rotation at Node No. 11 was less than the hinge rotation at Node No. 17. The hinge rotation at Node No. 17 was, as expected, smaller with two hinges than with only one hinge. 1 Three additional computer runs were made. Each of these made the stiffness of one of the support plates equal to zero, representing failure of one of the three aupport plates. In all three ca9es, the tube bending moments were much lower than the bending moments calculsted "cr the earlier model where all three support plates loaded the t abe. The cooled reheater helical tube bundle will experience local plaatic straining of les* than 0 3%, provided the tube support atruct 2re
- doe
- not auf fer creep collapse. If one of the three tube Support struc-l tures fail, causing the helical bundle to go out-o#-roJnd, the tube Page 49
909190 / A plastic strains will be less than one quarter of those plastic strains without the atructure collapsing. The pla9 tic atrain can now be com-pared to the expected minimum end-of-life strain at rupture, whi0h la 85 (see Appendix B). The calculated tube strain of 0.3% is muen less than the expected rupture strain for this one-time event. Therefore, the ] helical tube bundle is expected to remain a viable preaaure boundary under restrained thermal growth of the helical tube bundle during the firewater cooldoc.n event.
- 6. THERMAL SHOCK OF REHEATER TUBE BENDS
) It was deemed prudent to evaluate the effect of the thermal ahock-induced surf ace strains at the reheater tube bends. These amall-radius tube bends were formed by cold-working the Sanicro 31 material. Thia cold-working caused the material to be less ductile than the as-received tubing. Therefore, the material in these bencs may be more susceptible to cracking under rapid high strain loadings, even 1* of a thermal nature. This evaluation used the conservative hand calculation method of Ref. 11. The tubes were modeled as cylinders who9e outer surface waa assumed, for conservatia=, to be insulated. For the actual tube, the atagnant helium acts as an insulator of the outer tube sur* ace until I I convective flow o' the helium becomes establiahed. The "irewater temperature at the reheater tubes " rom Re". I will be 30 F. The water side Reynold's number waa calculated to be 5.55E04, and tne correaponding film coefficient was calculated to be 1550 Bta/ h-ft2 F. The tube dimensions are 1.125 in o.d. and 0.140 in, wall thickness. The material 19 Sanicro 31 (Alloy 800 Grade 2) which haa the following properties. Thermal conductivity k which 1
- almoat linear between 80*F and 1200of is taken " rom Re'. 12.
k = 81 3 Btu in./f t -h F @ 80 F and k - 152 Btu in./ft 2 -h F@ 1200*F Page 50
909190/A 1 Density - 0.283 lb/in.2 i Specific heat - 0.11 Btu /lb 'F. The coefficient of mean thermal expansion, as shown in Ref. 2, is not linear over the range 80*F to 1200*F. A value of thermal expansion will be determined based on the calculated tube wall temperature after the max has been calculated. surface temperature at at The stagnation temperature of tne reheater af ter 1.5 h and prior to introduction of the feedwater is assumed to be 1200*F (Ref. 1). The above values are used in the following calculation. See Ref. 11 for nomenclature used in this section of the docur: lent. Reheater tube o.d. = 1.125 in. Wall thickness - 0.140 in, r - 0.5625 in. e r - 1 0.4225 in. R - 0.7511 - r /r 1 e Use R - 0.8 as closest value charted in Ref. 11 Max. B -h.r/k- 80(0.%2O 81 3 Inaide sur" ace B 1 - 10.7 @ 100oF Min. B - 1550 (0.5625) i 152 Outside aurface B g - 5.7 @ 1200 F . Page 51
l 909190/ A It has been conservatively assumed that the strain producing at is the difference between the temperatures at the inner and outer tube sur- , faces, instead of the difference between the inner surface temperature and the mean tube temperature. 0 T = t t where t = 1200*F t,
= 80*F t = 1200 - 1120T .
Values of T are obtained from chart 36 of Ref. 11. and are tabulated in Table 6.1. This shows that the maximum temperature difference at acrosa the tube wall is 549*F. As this is greater than the temperature differ-ence between inner wall temperature and mean wall temperature, it will be used for conservatism. Assuming fully plastic surface deformation, i.e., Poisson's ratio (v) = 0.50, then the inner surf ace strain (c) is given by:
, , a at 2 (1 v) c - a at
~
a - 9.00 x 10 ' in./in.*F @ 573*F (from Ref. 2) c - 549 (100%) 8.68 x 10 tensile c = 0.49% 9 320*F at the inner sur# ace. A tensile surface atrain of approximately of 0.49% will be induced locally on the inner mur* ace of the reheater tube due to thermal ahock Page 52
909190 /A i TABLE 6.1 TUBE WALL TEMPERATURE DETERMINATION
- For R = c1 /r, = 0.8 F T t T D AT o 1 1 e e (B1 -10.6) (B-5.7) 1 0.001 0.275 892 0.0 1200 308 0.002 0.35 808 0.0 1200 392 0.005 0.47 674 0.02 1178 504 0.010 0.56 573 0.07 1122 549 0.020 0.67 450 0.20 976 526 0.030 0.75 360 0 31 853 493 i
0.050 0.85 248 0.49 651 403 0.100 0.96 125 0.77 338 213
- From Chart 36 or Ret. 11.
I l Page 53
i i 909190 / A loading. The minimum expected strain at rupture for Sanicro 31 cold-worked material (i.e., small-radius tube bends), is 8.0% (see Appendix B). This shows that tube cracking at the tube's most critical region will not occur from a single thermal ahock of the hot reheater tube bends with the cold firewater.
- 7. CONCLUSIONS The results of this structural evaluation of the most critical regions of the steam generator during a single cycle cooldown from power using firewater in one reheater module has shown the following resulta.
- 1. A single cooldown from power, using firewater in one reheater module, is not likely to cause cracking o' the cold-worked reheater tube bends. This appears to be independent of the power level at the initiation of the event.
- 2. The large temperature difference between the 80*F water and the high-temperature helium does not cause excessive plaatic straining of the reheater helical tube. The power level assumed for this evaluation was 50%.
3 The external pressure, combined with the hot temperatare of tne helium, leads to creep buckling of the tubes unless the temperature is limited. Assuming that the 35% power level helium pressure-time history 9hown in Fig. 4.6 does not change with change of temperature, and also assuming that the 9hape of the 35% power level temperature-time h19 tory shown in Fig. 4.7 does not change with change of peak temperature, then the maximum allowable helium temperature, including hot streak and hot module effect that occur at the bottom of the steam generator, must be limited to 1350*F during thia one-time event. This allows a higher temperature than the 1330*F shown c or 35% power level. Page 54
909! 90 /A
- 8. REFERENCES
- 1. Potter, et al., " Firewater Cooldown Uming the Reheater (1.5-Hour Delay) " GA Document No. 909113, October 1986.
i-
- 2. ASME B&PV Code, Section III, and Code Case N47.
t 3 Krishan, M., GA internal memo 306-4426, "BUCKLi. User's Manual," July 1980. l l
- 4. Pan, Y. S., " Creep Buckling of Thin-Walled Circular Shells Subject to Radial Pressure of Thermal Gradients," J. of Applied Mech. ,
l March 1971, pp. 206-216. I 1
- 5. Lindgren, J. , " Materials Property Input for FSV-HTGR Steam i
Generator Firewater Cooldown Analysis," GA memo MC:114:JRL:86, dated October 29, 1986. (See Appendix B of this document.)
- 6. GA Document CGA-D-41 Issue A, " Quality Assurance Procedure "cr the Colling and Bending Quali'ication," March 4, 1970,
- 7. Timoshenko, S. P. , and J. M. Gere, Theory of Elastic Stab.lity, McGraw-Hill, New York, 1961, pp. 289-297.
- 8. Faupel, J. H., Engineering De9ign, John *diley & Sons, New York, 1964, p. 248.
- 9. Johnson, J. J. , GA Report GAA-14006 DC-77, MODSAP Publication, June 1976.
- 10. Neal, B. G. , The Plaatic Methoda o* Structural Analyais, Chapman &
Hill, Ltd., New York, 1963, p. 27. 4 Page 55
909190/A
- 11. Schneider, P. J., Temperature Response Charts, John Wiley & Sons, ,
Inc., New York, 1963, p. 94
- 12. Wattier, J. B. , GA Report GAD 15431. " Preliminary Assesement of Structural Materials for Use at Temperatures in the Range of 1200*F to 2000*F," February 1980.
- 13. Perceira, N. G., "2 1/4Cr-1Mo Low Alloy Steel Materials Data Book,"
GA Report GAD 13817, January 1976.
- 14. Smith, G. V., " Evaluation of the Elevated Temperature Tensile and Creep-Rupture Properties of 1/2Cr-Mo,1Cr-1/2Mo, and 1 1/4Cr-1Mo-Si Steels," Material Properties Council Report DS50. ASTM published September 1973
- 9. CALCULATION REVIEW REPORT The Calculation Review Report is discussed in the following page.
Page 56
G A 15431R EV.11/803 CALCULATION REVIEW REPORT 90g190f3 TITLE: APPROVAL LEVEL QAL LEVEL DISCIPLIN E SYSTEM 00C. TYPE PROJECT DOCUMENT NO. ISSUE NO.. LTR INDEPENDENT REVIEWER: NAME / ORGANIZATION ' A - REVIEWER SELECTION APPROVAL: BR MGR - ' OATE # ' b
./ / /
REVIEW METHOD: YES NO ERROR DETECTED ARITHMETIC CHECK ' LOGIC CHECK # ALTERNATE METHOD USED I E N" '"' " ' ) SPOT CHECK PERFORMED
' b'I COMPUTER PROGRAM USED ' -
REMARKS: (ATTACH LIST OF DOCUMENTS USED IN REVIEW) he' .'J,9, !C . I ' e ', e manrM C CALCULATIONS FOUND TO BE VAll0 AND CONCLUSIONS TO BE CORRECT:
, ! ' ~', ', .(
INDEPENDENT REVIEWER ' ' OATE 6 ' ,' ',', _0, SIGNATURE
CA1543(REV.11/801 CALCULATION REVIEW REPORT TITLE: EFFECT OF FIREWATER C00LDOWN USING REHEATER ON STEAM APPROVAL LEVEL - GENERATOR STRUCTURAL INTEGRITY QAL LEVEL Q1 DISCIPLINE SYSTEM 00C. TYPE PROJECT 00CUMENT NO. ISSUE N01LTR. M 22 CFL 1900 909190 (CN5647) A INDEPENDENT REVIEWER: J. S. (Steve) Kaufman NAME ! ORGANIZATION Mechanical Engineeri m o REVIEWER SELECTION APPROVAL: BR MGR
/ DATE 9 Dec. 86
./
REVIEW METHOD: YES NO ERROR DETECTED ARITHMETIC CHECK
- LOGIC CHECK / ~
ALTERNATE METHOD USED SPOT CHECK PERFORMED v
/
COMPUTER PROGRAM USED '! REMARKS: (ATTACH LIST OF DOCUMENTS USED IN REVIEW) L~ c m p .s e t . .
- n. r!. +,. e fla/ usui w, y o ypj o - r:'<.
CALCULATIONS FOUND TO BE VALID AND CONCLUSIONS TO BE CORRECT: INDEPENDENT REVIEWER .
,f /
SIGNATUR E DATE /1 !/ M'
/
Page 58 l _ _ _ . _ . , _ _ , . - - - - - - _ _ _ - . - - - - - - - - - - - - - - - -
l 909190/A APPENDIX A
. THERMAL DATA 0F STEAM GENERATOR REHEATER SUPPORT PLATE AND TUBE DURING FIREWATER C00LDOWN l
1 Page A-1
909190 /A nmusur. maassRausmat GA 1976 Pas: 1900 EBGE: D. P. Carosella j*),g* C. === REFER E
._ ________ __.__..._--____- .._ ------.. -SE:DPC:208:86 . .
W: A. Shenay Dmt: 10/28/86 les: 2970.106 narIsun 10/30/86 SBJE2: PSV Stem Geraratcc Support Plate and Tube Tenperature During Firerater 0:cidown SUltSRY Data fran the SUPEREEAT code analysis of firewater coolcbwn fran 50% power on the reheater (Run SS372) was used to determine the tube support plate, tube axial temperature gofile and tube wall radial temperature gradients. 2ese results, whickt are gesented in Tables 1 and 2, can be used for the evaluation of "bearhug" stress in the reheater kundle. MALYSIS Re SUPERHEAT analysis includes a calculation of the gas temperature, the water j temperature and the inside, outside, and mean tube wall temperatures. Addi-tiorally, this analysis calculates the gas and water side film coefficients. - ne mean tube taperature and the tanprature gradients were tatulated in Table
- 1. Se gas and water tanperatures and film coefficients were used in a " hand calculation" to evaluate the support plate tempratures.
Se hand calculatioca made use of a one-D mndaction modal for radial enchetion in a disk. Se disk was centered at the center of a tube where it passes through the suppet plate. Se inside radius of the disk is at the surface of the clearance hole for the tube, the outside radius of the disk was based on a mathanatical combimtion of the transverse and longitudinal pitches. Se model accomted for the heat transfer resistance dae to the water side film coefficient, the tube wall-condactivity, the total radial resistance between the tube wall and the plate including all of the radial gags, and the radial conchetion in the plate. Beat into the plate was based on the gas side film coefficient and the temperature difference between the gas and plate. 21s calculation method is acre fully described in the attached calculation sheets. To facilitate this analysis the hand calculation method was included in a simple omnputer progra (SG* PLATE) . 2e results of that gogra are included in rm S3067 Rese results were verified by comparison to hand calculations included in the attached calculation sheets. PMa A-2 _ . _ - . - . _ _ _ , . - _ _ . _ _ _ _- . . _ _ _ _ - - ~ . _ . _ . _ _ . _ ~ . . _ . _ - - - _ - .
909190 /A A. Shenay 10/28/86 RI5F3 SED: IQ/3y86
'!he data presented in this meno and the attached calculation sheets represent a mnearvative aanbiration of support plate / sleeve and wedge / tube geanetry. Siis conservative configuration results in supprt plate tanperatures which are about 150F too hot at the top of the supprt plate.
~ ~
IPC:cc i cc: J. Kennedy A. IAwis R. Potter R. Ryder EDF P3ge A-3
909190 /A Table 1 Reheater Cboldcwn frczn 50% Pcwer Reheater 'Ibbe Wall 'D!mperature and Radial Gradient Distance frczn 'Ibp of Bmdle Mean Tube Temp. Radial Gradient (inches) 0F . OF 0.59 228 128 1.76 217 118 2.94 207 110 4.12 197 102 5.29 188 95 6.47 180 88 7.65 173 83 8.82 165 77 10.00 159 72 11.17 153 67 12.35 147 - 62 13.52 142 58 14.70 137 54 15.90 132 51 17.06 128 48 Pse A-4
909190/A
'mble 2 Reheater Coolcbwn frcm 50% Power Support Plate Temperature Distance frca Top of Btmdle Temp. l (irxtes) _ _ _ . . . . ..
oF l 0.00 47 3 1.18 449 2.35 42 5 3.53 403 4.70 383 5.88 364 7.06 3 46 8.23 330 9.41 314 10.58 299 11.76 285 12.94 273 3 14.11 26 1 15.29 249 16.46 23 9 17.64 229 Y Pag 9 A-5 i
c 2 m.v. . 2, CALCULATION SHEET gagigafa
~
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, , , , , , , , , , CALCULATION SHEET 909190/A CALCULATIONS FO R SUPPOW.T PLAT & TilH PEllATURS EQUIP. N O.
PROJ. j g g CALC.NO.FF* MI 20); ?( PAGE $ OF PREPARE 0 BY DATE REF. DOCUMENTS:
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CALCULATIONS FOR SoPPo g T PLATE Ten PER.ATUfLE EQUIP. N O. PROJ. 190o CALC.n0. $ 5
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u 2= w..v. u.2, CALCULATION SHEET 909190/A 90990(2r PL ATE TEtuFEEATUR.E EQUIP. NO. PROJ. g 9g CALC.NO.gg , pg / 203,* % PAGE 3 0F PREPARED BY CHECKED BY g { OATE DATE jgf p / g 4 R EF. 00 CUM ENTS: l 1 2 Fw 0 4 3 T ' 2 U% \ 2 ((, T. M )
.jt_ g g
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. g4 77g W RI = 0.7448 15 17 bt FMb
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- o. 3.. m .v. m,3, CALCULATION SHEET 909190 /A CALCULATIONS FOR 90FF0t.T 2 pt A7 g Tsaipy1ATURE EQUIP. NO. PROJ.
l339 l CALC.N0gg,,D#< ;2 d."FG PAGE / / OF // PREPARED BY g, { DATE 10 k l g REF. DOCUMENTS: CHECKED BY OATE ) Eh ?y 6 8 h 4
' O; -
6 Qt ki ( VH R.0 = -6IG./7 ( .47/ 2) ,
)
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] C) =
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" I
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909190/A APPENDIX B MATERIALS PROPERTY INPUT FOR FSV-HTGR STEAM GENERATOR FIREWATER C00LD0'4N ANALYSIS
)
1 1 Page B-1
909190/A 3 '
'l
) IN REPLY PROM J.R. Lindgre /J.B. Wattler REFEB 10 M C:114:JEL:36 10 R.H. Ryder/ Alan Lewis Project 2970106
( DM October 29, 1986 I SUBJECT Naterials Property Input for FSV-H'ItiR Steam Generator Emergency Cooldown Analysis The following data on materials properties were provided for creep collapse, thermal shock and bearhus analyses: l
- 1. GEEP COLLAPSE ANALYSIS 1.1 Creen Eaustions 1.1.1 T-2 (1/2 Cr - 1/2 Mo steel)
.$7492t.1x C$ G:gr9 4fCDh*2 I 5 *.
U 'U U Los t = g lo3 o + 2
+U log s, + _4 +U S 3 II' T .T . T
/YC Dh'l- bT To TMa~.M /wMwM Cf.nep itp ra i.t .*
B Log 4 ,= B +1+B..2 log o (2) T T
=
1.989551E+01 - 1 .34017 6549.50580 gog , T T __ L , e,* to Q = b' t Log a = 103 t + log i, Log t = Log - log i, I Substituting for log i ,Eq. (2) i Pu+ B-2
f 9093.90/A B B Los t = - log a + log e, - d - B 0 T T i 3 equating coeff in Eqs. (1) and gives U = -B 2 U =0 2 U 3
=1 U4 = -B 3
U S"~0 i, = %/hr t = time hrs e, = strain (%)
'T*'r-
- f* N f* O a = kai KUr*1UltL' PlObtt. FI T ~T*c .'b.i f$ ..
los t =B 0 1 2 3 8'
= 3.25299sE+01 - 0.01590 T - 0.07424 a - 0.00165 T log a Tertiary = 0.525 t e t ( 3000 he, temperature 90010F i 1400 e in ksi a > 0, a < min ultimate tensile strength at temperature T in A
1.1.2 T-22 (2-1/4 Cr - 1 No steel) " a 'u u Los t = d los a + [ +u 3 log s, + d + u S T T T P1de B-3
909190/A where t is time (hours), a is stress (ksi), se is creep strain (%), and T is absolute temperature (OF + 460) . Power Law Coefficients Complete Power Law
- Limited Power Law
- Annealed Annealed Alloy 800H 2-1/4 CR-1 MO 2-1/4 CR-1 MO (for 1200-1800*F)
U 1 -1.22 x 10 4 -1.126 x 10 4 -1.21626 x 10 4 U 2 -1.169 x 10 3 1.033 x 10 3 1.07849 x 10 4 U 2.125 3 0.7232 0.471666 U 4 5.64 x 10 4 4.913 x 10 4 5.14166 x 104 0 -25.49 5 -21.52 -19.0780 i a 0.145 gog , 0.116 900 - 1200*F 1000 - 1200'F Best fit l The time is limited by initiation of tertiary creep, t 3, can be esti-mated from the rupture life, t,, by t 3 = 0.525 t, , o>0, a < min ultimate tensile strength at temperature T where log t, = C, - 10.307 log o - 0.03691T + 0.004638T log a ( from NSMH) . N07E: a = MPa, T = *K (CK /VSMH fossran oxi Y. The lot constant C, had an overall average value of 47.253, with an average value of 47.395 for annealed material only. Pige 9-4
909190/A 1.1.3 Alloy 8005
'n
~
s u los t = d log a T
$+u 3 los e, + d + u$
T T
,-12162.6 g,g , , *
1078.49 + 0.471666 l o s e, + - 19.0780 T T T Log (t, + 3) = B0 + Bg /T + B2+B3 log a
= -17.0212 + *
+ 0.785712 - 12379.81 los e T
t, = t,,,g,,, in hours T = 'F + 460. 'F = 1200* to 2000'F a = kai, o > 0, e ( ultimate tensile strength at temperature T B0 = -17.0212 B = 48158.12 B 2 = 0.78572 B 3 = -12379.81 0.994 0.2% t3 = 0.385t r where t 3 = time to tertiary creep in hours tog &,= B +1+2 . log a T T
= -17.0212 + "* + 0.78572 los a T T Paga B-5
l 909190/A 1 1.1.4 Alloy 800, Grade 1 1 2 Los t, = B0+T~~+~~1ga 20 to 1000 hours T
= -1.668672E+01 + *' ' - 0621.82464 g,s ,
T T-t, = time to rupture in hours t 3 = time to tertiary in hours = 0.385't9 994 T = 'F + 460, 'F = 1200* to 2000'F a = ksi, a > 0, a < ultimate tensile strength at temperature T B Los i * =B + _I + B_log _ a T T
= 2.344083E+01 60398.73818 ,12993.75758 g, ,
T T i, = minimum creep rate %/hr 1.2 Youan's Modulus See attached sheet A-I for T-2 (1/2 Cr - 1/2 Mo), A-I, A-III A-IV for T-22 (2-1/4 Cr - 1 Mo). 1.3 Poisson's Ratio l See attached sheet A-V for T-22 (2-1/4 Cr - 1 Mo).
#e suggest that values for T-22 be used also for T-2 for which values were not found.
P389 B-6
l 909190/A ) 1.4 Yield Stronath See Figure 1 and la for yield strength vs temperature for T-2 and Fig. 2 and 2a for yield strength vs temperature for T-22.
- 2. T1tERMAL SHOCK ANALYSIS 2.1 Lowest tensile ductility for Alloy 800H tubing in temperature range RT l
to 600*F. taking into account thermal aging environment, and cold work. The effects of thermal aging are the most significant at 1100'F ( 600'CI ' and those effects on room temperature ductilitr appear to saturate at about 40,000 hours so that at the end of life the minimum RT tensile ductility would be 8% total elongation. This is true for both solution annealed and cold worked Alloy 800H. The tensile ductility of cold worked material initially is less than half of that of solution annealed material, but upon thermal aging at temperatures of about 1100'F for 40,000 hours the ductilities of solution annealed and cold worked Alloy 800H are essentially the same (per discussions with Jude Foulds of GA and J.M. Martin of Inco International, Inc.) . {
- 3. BEARBUG ANALYSIS
\
The lowest RT to 600'F tensile ductility for the coiled part of Alloy 800H tube bundle (i.e. , not tube bends) taking into account thermal aging and environment is 8% total elongation. This minimum is reached due to thermal aging effects as discussed in Section 3 above. JRL:MDR cc: J.R. Foulds W.R. Johnson D.I. Roberts P1ge B-?
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e fy Q Firure 1. Effect of temterature on yield strencth, tensile w.m = 5 strencth, Oreer strength (0.01 4 per 1000 hours,. and ruuture .,.. iqi W strenr*n (100,000 hours) of 1/2 Or - 1/2 Me steel. Yield and = i - tensile strenr*hs are ad justed to 30 and 60 ksi at ?! ?. Creer _M =g and rupture strenetns are averates of available data. m
=
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909190/A r- == __ g.--_]
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l l! _-- t - Figure la. Variation of yield strength of 1/2 Cr - 1/2 Mo steel l with temperature. 75 l Pag- 9-9
909190j; 3
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--g
; Figure 2.-
g Effect of temperature on yield strength, tensile
# strength, creep strength (0.01 % per 1000 hours) and rupture . ..
strength (100,000 hours) of annealed 2 1/4 Cr - 1 Mo steel in -[g wrought er cast form. Yield and tensile strengths nave been adjusted to 3C and 60 kai at 75 F. Creep and rupture strengtns y are averages of available data. .. T ' f h ',
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l
. , oscaussa 1sn FERROUS ALLOYS [ -[
x s~ -- 909190/A
;.' 100 l
3,- T-t (A alt ri o T-1 Fe ! 2DtFLATE O.15 C
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[ 0.25 Si t' [b. sw Cute Rur1 RE N """ ' N ,: i e : R 10tsr. a i j A , oo ,, e t0.000 uCUR i f f O NO $ 3 700 800 900 1000 1100 Y"l go
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- TEMrERATcRE -r E45
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r:2. 3.=1 Cute acrwRE TuNcTs or T-1 STEEL AT VARICUS TEM-PERATURES. (13) b A '
. AsrM TEsrs et' 81, -m -m -100 ei 100
- TExp- r s
<s 0 220s
.i- ~ so m .
cw I T-t
.g . , , yrc. 3.ssf23 EFFECT OF TEMPERATURE ON P1.ANE
- s. - STRAIN FRACTURE TOUGHNESS OF 2 Di m 3/4na sus F , a 121 m Fw = 112 tsa (31. table 2. Ag.21) y-a ..a
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YO TOUGletESS OF PLATE. 1 : 2 r_
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Page B-12 nos 11
k[ 909190/A PRCPERTY CCDE 2111 II PAGE NUCLEAR SYSTEMS MATERIALS HANDBOOK " in?" PART I - STRUCTURAL GROUP 2 - LOW ALLOY STEELS SECTION 2 - 2 1/a Cr-1 Mo MATERIALS REVISION: 1 5 l-77 ] PREVIOUS RE ISION: 9-30-76 YOUNG'S MODULUS . 7A ct E I - 3 7A7:C YOUNG'! -COULU:
*
- S I U h !T O (U7 UNIT! IM 3:F3NTHf5E3)***'JS UNI *3 (S! U*,ITS IN 840ENTHEIPS)-*
T EMc . s7PE!! = 7E"o. ST=riS 3:0 C (F) !! 12aSCALO *ILLION 85I
- 7EG F ( *^ 3 MIL LION CSI GIO:ca;;*Ly
[ 24( 751 2.:33r*CZ ( 3.C21F+01)
- 75( 243 3.021E*C1 ( 2.J33EC '1 E 25( 771 ?. S 2 e.r 2 ( 3. 19F**1)
- LCC( 3d3 2 917E+01 ( 2. 6;F-:21
?C( 1221 2. :53 E*G 2 ( 2.97ai+01)
- 15C ( 66) 2.956E*11 ( 2. l TS E *C1 7?( 167) 2.: ;r 02 ( 2.94 g*013
- 20:( 9!3 2 9225*41 ( 2. 15E*021 10 . ( 2121 ?. C13 c *0 2 ( Z . 910i +C 11
- 290( L211 2.395E.01 ( 1.?95E+ 21 127( 2?71 1.19 4 r +0.dk ( Z.892r+'13
- 3 :( 149) 2 474#*71 ( 1.151?+001 150( 3C2) 1 331 E*G Z ( 2.871'+:13
- 35C( 177) 2.5ESE+C1 ( 1.16 9 e
- C ' 1 17?( 3*73 1. 969 E *0 2 ( 2.*T6E*013
- 4:0( 2* 43 2 439E*ft ( 1.19'E*0?)
Zart 3c21 1 979 E+3 2 ( 2.
- L i'. *31)
- 4?C.C 2271 2.323E+C1 ( 1.146E*021 22?( 4J71 1.?49E*C2 ( Z . S2 7E *'.11
- 51C't'26G) 2.aCSE+G1 ( 1.135E+C21 25"( 4223 1 939 E+12- ( 2.812:+;1)
- 5?L( 2*9) 2.7 86 E
- 01 ( 1.321E*C?3 27?( F27) 1. 72 9 E+0 2 ( 2.786:*e13
- 60C( I1Al 2.763E*i1 ( 1.105E+021 30Ct ? 72 1 1. 914 E+0 2 ( 2.777c+C1)
- Sitt 3421 2.734E+C1 ( 1.19?C*Ct1 325( 6171 1. 4 99 :+4 2 ( 2.7!4E**.11
- 7::( 371) 2.498'+01 ( 1.9 ACE +C2)
- 3?C( 6621 1 379 E+0 2 ( 2 7265+J1)
- 7?C( 3991. 2.6 53 E + C 1 ( t.12SE*CII
'37!( 7C71 1.d?6E+CZ ( Z.692E*C1)
- aCC( s271k 2.595E+01 ( 1.792E+021 -"'
- 4. c t 7?21 1 323 E+0 2 ( 2. 6 F 1 ~ *013 *
$5C( 454) 2.5.TZ E *C 1 ( 1.746E*c2.1 425( 787) 1. 794 E *C Z ( 2.6'2e**11
- 9' f ( 642) 2.4 E3 E *** 1 (
- 1. 411 '
- C 11 45C( 8421 1. 754 E+0 2 ( Z . ?baE +b il
- 95CC 51:1 2.319F+G1 ( t.62CE*C21 47e( Ae71 t. 7:6 E+C 2 ( Z.477E*lil
- 1J00( ?343 2.2 43 E *C 1 ( 1. 7Y L E *R11
?C"( 9321 1.651E*C2 ( 2.299E*31)
- 105G( 566) 2.121E+Gt ( 1 463E*C01 32TI 9773 1. 33 7 E +C 2 ( 2.3: 2 E 691)
- 11*0( ???) 1 975?*71 ( 1.150c*C21 3 ctic221 1. ?i3 E+0 2 ( 2.195E.C1)
- 115;( 6211 1.3055 11 ( 1.246E*00)
F7?(10F7) 1. 43C E *G 2 ( 2.C743*C1)
- 12C3( 6 4'23 1.613 E *l i ( 1.115E *G 2 3 500(11121 1. 335 E+G 2 ( 1.937i+41)
- 52*(11E7) 1. 229 E+* 2 ( 1.792c* 1)
- S49(12C33 1.115E+02 ( 1 615E-;11
- T HE VALUES A9CVE WE " : LCULA753 F,.1.* THE FCLL] WING ACLYN]w AL FCUATICN: 'N Yu a (* 3 *4 (1) X *. . . A ( Fi x** , wHi3 E t y = S TR g S S , ,1 PS* ys7EMP.,CEG =
1(31= 3.1173*:E* 1 . -"' A(113 -1.354000E-;Z a ( 21 = 2. 7 C T 2 6' ' E .' S A(T1= -2.;c770GE-c6 TO C3NVERT FRCM U! TC TI UNI 73. .=ULTIoLY -?- 3Y 6. 8947? ?E+C 1 s 90?Ft THd E10a72QN FAPG2 I3 2.-643 OgG C (73-ggg; QuG 2). N 4......... ........................-....... ... ............. J 8/1-50 Page B-13
) w 77 "'E.
909190 /A
;TY CCDE . 2111 PAGE 1.2 l NUCLEAR SYSTEMS MATERIALS HANDBOOK " ' - ESJ" PART I - STRUCTURAL GROUP 2 - LOW ALLOY STEELS MATERIALS SECTION 2 - 2 1/A Cr-1 Mo REVISION: 0, 5-1-77 YOUNG'S MODULUS (Cont'd)
TABLE II - OYNAMIC YOUNG 'S MODULUS
** S I UN IT S (US UNITS IN P19ENTH6SESl ***US UNITS (SI UNITS IN P A PEN THESES 3 **
TEMP. STRMSS
- TEMP. STRESS DEG C (F) GIGAPASCALS MILLION PSI
- DEG F (C) MIL LI CN PSI GCGAPASCALS 24( 751 2 162E*02 ( 3.135E+Cil
- 75( 2 41 3 135E+01 ( 2.162E+02%
SG( 1221 2 176E+02 ( 3 0978+01)
- 122( 38) 3.115E+C1 ic e" ( 2121 ( 2.147 E + 0 21'
- 2. 3 9C E+E 2 ( 3.J 31E +01) + 23C f 9 31 3.040E+.1 ( 2.c96E+021
.15ft 3C2) 2.;49E+P2 ( Z.972E*41)
- 330( 149) 2 977E++Il ( 2.05GE+02T 220( 39?) 2. C11E *0 2 ( 2.917E+cil
- 40C ( 2c el 2 912E+31 ( Z.335E+02) 25 C ( 4d21 1 97pE+G 2 ( 2.e65E+ ell
- 9 )C ( 260) 2.855E+C1 ( 1.?68E+02) 3C: ( FT2) 1.941E+02 ( 2.315E+.11
- 6 C t. ( Tis) 2 799E+01 ( 1 930E+w2) 35C( 6621 1 916E+02 ( 2.7'34E +01)
- TUC( 17 11 4GCt 7521 2 . 74? E* 4L-. ( 1.390E+02) 1 469E+02 ( 2.711F+C1)
- 3.0 G ( 427) 2.631E*C1 ( t.a49E+02) 450( 5421 1. F3')E+0 2 ( 2 654E+tt)
- 1;9 )G ( -4 2) 2.614E+11 ( 1. a O 3E +0 2) 501( 932) 1.737E+C2' ( Z.592E+C1)
- 1000( 535) 55C(1022) 2 . 5-l E+ c t ( 1.75tE+02) 6 00 t111?)
- 1. 739 E +0 2 ( 2 522E+cil +* 113C t $9 3) 2.a5 4 E+ 01 ( 1.692E+02)
, 1 696E+02 ( 2.443E+G1) 12cc ( s49) 2.356E+ci ( t . 624E +0 2 )
650(1202) 1.623E+02 ( 2 394E+011
- 132C f 7C 41 2.242E+01
~,,.~~ 720(1292) 1. 53 2E+0 2
( t.546E+02) ( 2.252E+Cil
- 14JC ( 764) 2.113 2+ 61 ( 1.455E+02) 7?C(1382) 1.472E+r2 786(18531 1.404E+02 ( 2.135E+til
- 1453 ( k3 b di
- 2. C 37 E+ C1 ( 1.k44E+c2)
( 2.036E+R11 *
..................................................................... 1....
THE V ALUES 420VE WEFE C ALO(: LATED FROM THE FCLLOWING POLYNOMIA L EQUATION 8 E= A ( TF) * (1+ 3 (T-T F) +C ( T-TF1 ** ?+ 0( T- TFl '*I} , WHEREI Es OYNANIC YOUNG ' S MO CULUS, MILLION PSI. T= DEGREES FAHRENHEIT
?
T Ws 7= OE7 F A(TF)= 31 35E6 l B = -2. 46-4 ( +/ 2.0 M E-t. ) Ca 1.8E-7 ( +/-G . 3 2E- 7 ) Oz -1.2tF-10 (+/-C.28E-1") s
' (STANPARD DEVIATIONS OF THE COEFFICIENTS ARE GIVEN IN DtRENTHESIS.1 TO CONVERT FROM US TO SI UNITS, MUL TIot.Y "Y" $Y 6.394757 NOTE! TPS 5,2UATION RANGE IS 2'*.7:12
. ...... .... . . ......... ggr 0 (7 J.14?c O
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L- l 11 NUCLEAR SYSTEMS M ATERI ALS II ANDDOOK voi. :-DESIGN DATA rin - neueruman. uanmu l caour : - inw Auor sucis l socrion a - a na er - i u. anymon .-i-n raevious uvmon -n-n YOUI'!G'S MODUI.US (STATIC AND DYNAMIC) TEMPERATURE . DEGREES FAllRENilEIT 200 400 600 800 1000
' ' ' 1200 1400 220 '
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q; ' l-~ 3 0 16 100. 200 300 400 500 600 700 800 TEMPERATURE . DEGREES CEl.SIUS
- y 4 ,,uc m , ucr i m us s .... ,,- . . . . . . ,,4 s : ,u. s e m -r m ,,... ,, m o , .. m m ,. ...a ..a. is e m -r u .1 ... ..... .., 7 -%.
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'y A,- 909190/1 RHi@g GENERAL ATOMIC COMPANY
, PART:
"/O uE '/
GROUP: I-STRUCil1RAL MAEIALS MATERI A(.: 1 2 LOW ALLOY u c.i.5 21/4 Cr - 1 Mo ; REVISION: -0 DATE: 11 1 75 Tm.E: PROPERTY CODE: 2110 : POISSON's RATIO (0YtlAMIC Att0 STATIC) l 3
- 1. Applicability 1.1 Product foms per saecifications lis ed beicw
- Seamless Tubes and Pipes SA 335-P22 SA213-T22 Plate SA 387-Gr22 Bars and Rods SA 335-F22a I 1.2 Annealed material anly
- 2. Grauhical Data: Graph an subsecuent page is a plot of ttm equation below.
The data points on which these equations; are based will.be incorecrated in a subsequent issue gf this propervj. l-
- 3. C nstitutive Equation
,i
~ Static V= 0.254 + 1.54 x 10 T-1.25 x 10~I T 2 (75 *Tdl200*F)
Dynamic V= 0.259 + 0.2 x 10'IT+7x10-8EN.y T - I x 10'II 2 T(75s Te 1450*F)
- 4. Data points (from ecuation above): '
'f, 5~ATIC Y' DY?!AMIC 17 2 STATIC v' OYtlAMIC V 75 .254 .259 700 .300 .233 100 .258 .250 750 .299 .235 150 .274 .251 800 .297 200
.237 ;-
.280 .252 850 .294 .290 250 .235 .253 900 .291 .292 200 .289 .255 950 .287 .294 350 .292 .257 1000 .282 .296 400 .295 .259 1050 .277 .298 450 .298 .271 1100 .271 .299 500 .299 .273 1150 .254 .301 5~0 .201 .275 1200 .257 .302 600 .301 .278 1200 .303
- 550
.301 .280 1400 .303
- The n-, analysis
. . s.. . c is based on data from material conforming to these certificati0"5-
..a
909190/A INTERNAL CORRESPONDENCE GA 1076 IN REPLY FRON: J. R. LindgrerI/J. B. Wattier PEFER TO TO: R. H. Ryder/ Alan Lewis I DATE:11/12/86 l
SUBJECT:
Addendum to " Materials Property Input for FSV-HTGR i S':eam Cenerator Emergency Cooldown Analysis" MAB:114:JRL:86, Dated October 29, 1986 At the request of W. R. Johnson, the references applicable to the materials input given in memorandum MAB:114:JRL:86 are given below.
- 1. Sterling S. 4
" Temperature-Dependent Power Law for Monotonic Creep" GA-A13027, March 1976 2.
Aerospace Structural Metals Handbook, AFML-TR-68-115 1979 Publication, Code 1103, page 11, Fig. 3.062. Modulus of Elasticity at Room and i
'levated Temperatures 1/2 Cr - 1/2 Mo T-1 steel.
3. Wattier, J. B. to W. G. Hardin "A Temperature Dependent Creep Equation l for Carbon Steel" 209:25:78, February 14, 1978.
- 4. Smith, G. V.
" Evaluation of the Elevated Temperature Tensile and Creep Rupture Properties of 1/2 Cr - 1/2 Mo, 1 Cr - 1/2 Mo, and 1 1/4 Cr - 1/2 Mo Si Steels" Prepared for the Materials Properties Council, ASTM DS50, 1973.
5. Smith, G. V. " Supplemental Report on the Elevated Temperature Properties of Chromium-Molybdenum Steels (on Evaluation of 2 1/4 Cr - 1 Mo Steel) Prepared for the Materials Properties council, DS6S2, 1971. 6. NRIM Creep data shee: No.-3 (1972) and No. 11- (1974) " Data Sheaets on -~ - the Elevated Temperature Properties of 2.25Cr-1Mo steel. 7. NSME Vol 1 Design Data,-Property code 2206 Part 1 Structured Materials, Group 2, Low Alloy Steels Section 2, 2 1/4 Cr - 1 Mo, Revision 1, Creep Equation Page 1.3, Young Modulus, pages 1.0, 1.1, 1.2; Revision 0 Property Code 2110, Piosson's Ratio. 8. Wattier, 1980. J. B. " Alloy 800H Creep Equation" GA Doc 906211 page 6, Aug. 9,
- 9. Wattier J. B. to D. I. Roberts " Determination of Maximum Allowable Primary Membrane Stress, So, for Tempetstures in The Creep Range for Alloy 800H" MAB:29:JBW:83 March 3, 1983.
- 10. Schill, T. V. to M. P. Ranson "Incoloy 800 Data Package" No.C-45,
, Huntington Alloys, Inc. Div. of Inco, March 24, 1981.
- 11. Telecon J. R. Lindgren, GA and J. M. Martin Inco " Minimum Residual Tensile Ductility after Thermal Aging Alloy 800 grade 1 and Creep 1
Page B-17
909190/A-Equations / Data for Alloy 800 grade 1 up to 1800 F," October 17, 1986.
- 12. Discussion J. R. Foulds and J. R. Lindgren GA " Minimum Residual Tensile 1 l
Ductility of Alloy 800/800H after Thermal Aging" October 16, 1986. l m-Page B-18
-o-- -
I '/ J) - i 8 NUCLEAR SYSTEMS MATER ALS HANDBOO( vou -= = .... p . _ ,_ ...o. ,,,,,, i,<,.o.....o._...o......,
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- o. -l l'i' l- 1 l ~ l'l ti-l-1 1-1 1 1 1-e e e i lj ggg, O 10 0 200 300 o 400 500 600 700 800 900 t000 y TEMPERATURE, DEGREES C S y *
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ELEVATED.TEWp(RATURE AND Roces. TEMPERATURE YtELD STRE88GTHS '^
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