ML20083C889
| ML20083C889 | |
| Person / Time | |
|---|---|
| Site: | Clinch River |
| Issue date: | 08/31/1974 |
| From: | Thomas K WESTINGHOUSE ELECTRIC COMPANY, DIV OF CBS CORP. |
| To: | |
| References | |
| WARD-D-0010, WARD-D-10, NUDOCS 8312270093 | |
| Download: ML20083C889 (109) | |
Text
{{#Wiki_filter:_ i
"" WARD-D-0010 j,I,I .
i
'l [LMFBR <
~a -
l t - H i Clinch River
; Breeder Reactor Plant Nuclear Island h
c SELECTION OF COOLANT-BOUNDARY ' MATERIALS FOR THE CLINCH RIVER ' ( l BREEDER REACTOR PLANT AUGUST 1974 s t\oo p' D i h* J. Prepared for the Project Management Corporation as part of the U.S. Atomic Energy Commission 1.igmd Metal Fast Breeder Reactor Demonstration Program Any Further Distribution by any Holder of this Document or of the l Data Therein to Third Parties Representing Foreign interest, Foreign Governments, Foreign Companies and Foreign Subsidiaries or Foreign Divisions of U.S Companies Should be Coordinated with the Director, Division of Reactor Research and Development, U.S. Atomic Energy Commission W Westinghouse Electric Corporation ADVANCED REACTDRS DIVISIDN ,p l , ,, ' MADISDN, PENNSY LV ANI A 15663 e l D DO 050 $37 1} A PDR
}
y 1 D This report was prepared by the Westinghouse Advanced Reactors Division as an account of work sponsored by Project Management Corporation and the United States I,overnment. lieither the United States nor the United States Atomic Energy Commission, nor Project Management Cor-7' poration, nor ary of their employees, nor any of their contractors, subcontractors, or their employees makes any warranty, expressed or 3 implied, or assumes any legal liability or responsibility for the o accuracy, completeness or usefulness of any information, apparatus, product or process disclosed, or represents that its use would not infringe privately owned rights. D 1 J
.y-g ; : c-
~
o "o
?
9
- g. .
b. SELECTION OF COOLANT-BOUNDARY MATERIALS FOR THE-CLINCH RIVER BREEDER REACTOR PLANT P. 500 I
?
1 es
. eb
' APPROVED:
/ ~/
) ,s
.y
- i. c . 4/
K. C. Thomas i, J4anager Materials Erigineering Westinghouse Electric Corporation Y- c Advanced Reactors Division P. 0. Box 158
- -Y Madison, Pennsylvania 15663 August, 1974
w
.o TABLE OF CONTENTS PAGE NO.
vii
SUMMARY
viii LIST'0F TABLES LIST OF FIGURES ix xii
' ACKNOWLEDGEMENTS
- 1. INTRODUCTION 1
- 2. PROPERTIES OF TYPES 304 AND 316 STAINLESS STEEL 3 2.1 Mechanical Properties 3 2.1.1 Short-Term Tensile Properties 3 2.1.1.1 Effects of Carbon and Nitrogen 3 2.1.2 Creep Properties 7 2.1.3 Fatigue Properties 9 2.2 Sensitization and Heat Treatment Effects 10 2.2.1 Measurement o' Sensitization 11 2.2.2 Occurrence of Sensitization in LMFBR Components 11 2.2.3 Effects of Sensitization 12 2.2.4 Comparison of Alloys for Resistance to Sensitization Effects 12 2.3 Stress-Corrosion Cracking 14 2.3.1 Effect of Environments on Chloride Stress-Corrosion Cracking 14 2.3.2 Effect of Temperature on Chloride Stress-Corrosion Cracking 15 I 2.3.3 Effect of pH Value on Chloride Stress-Corrosion Cracking 15 2.3.4 Effect of Alloy Content on Chloride Stress-Corrosion Crackin9 15 iv
_ . - - = . J A l 2.3.5 Effect of Structure on Chloride Stress-16 Corrosion Cracking 16-2.3.6 Stress Corrosion in Caustic Environments 17 2.4 Sodium Corrosion Rates 18 g
-2.5 Fabrication and Weldability 18
- 3. PROPERTIES OF FERRITIC 2-1/4 Cr-1 Mo STEEL 19 1
3.1 Mechanical Properties ~ 19 3.1.1 Short-Term Tensile Properties 20 3.1.2 Stress-Rupture Properties ~ 20 3.1.3 Fatigue Properties 20 3.2 Stress Corrosion Cracking 20 3.3 Sodium Corrosion Rates 21 3.4 Steam Corrosion Rates 22
': 3.5 Sodium-Water Reaction Effects 22 3.6 Fabrication'and Weldability 22 3.6.1 Hot Forming 23 3.6.2 Welding.
3.6.3 Welding of 2-1/4 Cr-1 Mo to Austenitic 23 Stainless Steels 27
- 4. INTERSTITIAL TRANSFER EFFECTS 27 k 4.1 Selection of Carbon Potential Values 28 4.2 Compensation ~for kItrogen-Transfer 28 4.3 CGlculation of Interstitial Gradients 29 4.3.1 Primary System Hot-Leg Components 30 4.3.2 Primary System Cold-Leg Components 30 4.3.3 Intermediate System Hot-Leg Components 30 4.3.4 Intermediate System Cold-Leg Components '
31 4.3.5 Intermediate System Ferritic Components
~
4.4 Comparison Between Low, Normal and High-Carbor t 31 Grades I 32
- 5. COST CONSIDERATIONS v
6.. GUIDELIf1ES FOR ALLOY SELECTIO'l . 33 6.1 Selection for Interstitial Transfer and Mechanical Behavior 33 1 6.2 Selection for Fabricability, Weldability and Sensitization 34 6.3 Selection for Stress-Corrosion Cracking Resis-tance 34 6.4 Tentative Recommendations for Material Selection 35
- 7. REFERENCES 36 TABLES 44 FIGURES 64 e
vi
v b S'JMMARY A study has been made of the pertinent material properties which will influence the selection of a particular grade of austenitic stainless steel in primary and intermediate system CRBRP components.
' Types 304 and 316 stainless steel were considered, together with their low and high-carbon derivations. General-guidelines were established which will enable the optimum choice of material to be made. A description of the properties of ferritic 2-1/4 Cr-1 Mo steel is also presented in order to define areas where it could possibly be used in place of stain-less steel. It should be noted that 2-1/4 Cr-1 fio steel has already been selected as the steam generator reference alloy, r 'M s
t. t vii
0. LIST OF TABLES _
- l. Carbon Contents for Austenitic Stainless Steel Product Forms.
- 2. Sy and Sm Valoes for Austenitic Stainless Steels.
- 3. Corrosives Reported to Induce Intergranular Corrosion in Sensitized I Austenitic Stainless Steel. i
} 4(a). _ Effect of Annealing Temperature on Reducticn of Stress in Austenit c Stainless Steel ) Effect of Annealing Temperature and Time on Reduction of Stress in 4(b).
Type 347 Pipe. Effect of Annealing Temperature on Relief of Stress in Type 316 4(c). Cold Drawn Tubing.
- 5. Results of Huey Tests on Several Heats of Nitrogen-Alloyed Stainless Steel.
- 6. Huey Test Data for AISI 304 and Jones and Laughlin 304-N (0.25 Nitrogen).
-7. Chemical Composition of Ferritic 2-1/4 Cr-1 Mo Steel.
- 8. Low Cycle Fatigue Data for 2-1/4 Cr-1 Mo Steel.
- 9. Test Conditions for Steam Corrosion Tests on 2-1/4 Cr-1 Mo.
- 10. Estimated Steam Corrosion Allowances for Ferritic 2-1/4 Cr-1 Mo Stee
~
- 11. Wastage Rates in Sodium / Water Reactions.
- 12. Coefficients of Linear Thermal Expansion for Type 304 Stain'less Steel and 2-1/4 Cr-1 Mo Steel .
- 13. Chemical Composition of Filler Metals for 2-1/4 Cr-1 Mo Welds.
14(a). Diffusion Depths for Hot-Leo Components.
.14(b). Diffusion Depths'for Cold-Leg Components.
- 15. Costs for Stainless Steel Plate.
- 16. Summary of Austenitic Stainless Steel Performance Under Lt1FBR Conditions.
17(a). Tentative Material Selections for Demonstration Plant Hot-Leg Components Tentative Material Selections for Demonstration Plant Cold-Leg 17(b). Components I L viii
v. 4 LIST 0F FIGURES _
- 1. Nitrogen /Carbot Ratios From Type 304 Stainless Steel Literature Survey.
- 2. Comparison Between _the Strengths and Ductilities of Solution-Treated Types 304 and 316 Stainless Steel Containing 0 Weight Percent (C + N).
i ?
- 3. Comparison Between the Strengths and Ductilities of Solution-Treated Types 304_ and 316 Stainless Steel Containing 0.04 Weight Percent (C + N).
- 4. Comparison Between the Strengths and Ductilities of Solution-Treated Types 304 and 316 Stainless Steel Containing 0.08 Weight Percent (C + N).
- 5. Comparison Between the Strengths and Ductilities of Solution-Treated Types 304 and 316 Stainless Steel Containing 0.12 Weight Percent (C + N).
- 6. Stress- Rupture for Austenitic Stainless Steel Bar'and Plate.
. .7. Minimum Creep Rates for Austenitic Stainless Steel Bar and Plate.
3
- 8. Effect of Interstitial Content on the 10 Hour Rupture Strength of
. Type 304 Stainless Steel Bar.
5
- 9. Effect of Interstitial Content on-the 10 Hour Rupture Strength of Type 304 Stainless Steel Bar.
3
- 10. Effect of Interstitial Content on the 10 Hour Rupture Strength of Type 316 Stainless Steel Bar.
5
- 11. Effect of Interstitial Content on the 10 Hour Rupture Strength of 1
Type 316-Stainless Steel Bar. J
- 12. Effect of Sodium Exposure on the Creep Rates of Type 316 Stainless a
Steel at 1325 F.
- 13. Onset of Sensitization in 18-8 Stainless Steel as a Function of Carbon as Determined by the Strauss Test.
ix
O 5
- 14. Effect of Molybdenum on the Time-Temperature-Sensitization Diagram of Type 304 (0.054% C) Stainless Steel.
- 15. Time-Temperature-Sensitization Diagram of Type 304 L Stainless Steel (Huey Test).
I
- 16. Time-Temperature-Sensitization Diagram of Type 316 L Stainless Steel (Huey Test).
- 17. Effect of MgCl Concentration 2
on Time to Failure.
- 18. Effect of Increasing the Nickel Content on the Susceptibility of Fe-Cr-Ni Wires in Boiling 42% MgCl2 '
- 19. Composite Curves Illustrating the Relative Stress Corrosion Cracking Resistance for Commercial Stainless Steels in Boiling 42 Percent Mag-nesium Chloride.
- 20. Effect of Carbon on Times to Failure of Stainless Steel in Boiling MgCl '
2
- 21. Effect of Heat Treatment on the Yield and Tensile Strengths of 2-1/4 Cr-1 Mo.
- 22. Effect of Heat Treatment on the Total Elongation of 2-1/4Cr-1 Mo.
- 23. Yield and Tensile Strengths of Low-Carbon 2-1/4 Cr-1 Mo Steels.
- 24. Summary of the Creep Rupture Properties of Annealed and Normalized and Tempered 2.25 Cr-1 Mo Materials.
- 25. Stress Rupture Curves for Laboratory Heats of 2-1/4 Cr-1 Mo Steels With Different Carbon Contents.
- 26. Effect of Carbon Content on the Rupture Strength of 2-1/4 Cr-1 Mo Steel at 1100 F (593 C).
X
e
~
t
- 27. Effect of Temperature and Oxygen Contamination on Corrosion Rate of 2-1/4 Cr-1 Mo Ferritic Steel .
- 28. Corrosion Rates for 2-1/4 Cc-1 Mo and Type 304 Stainless Steel in Liquid Sodium at 1200 F.
4
- 29. Corrosion of 2-1/4 Cr-1 Mo in Superheated Steam at 950, 1000 and 1100 F.
w'
- 30. ' Corrosion of 2-1/4 Cr-1 Mo in Superheated Steam at 1200 F.
31 '. Assembly of a low-Alloy Steel Pipe and a Stainless Steel Pipe that was Joined with Crack-Free Welds by use of a Nickel Alloy Filler Metal.
- 32. Interstitial Gradients in Primary System Hot-Leg Components.
- 33. Interstitial Gradients in Primary System Cold-Leg Components.
- 34. Interstitial Gradients in Intermediate System Hot-Leg Components.
- 35. Interstitial Gradients in Intermediate System Cold-Leg Components.
.i xi
9 ACKNOWLEDGEMENTS Acknowledgements are extended te S. J. Orbon for conducting the inter-stitial't'ransfer computer studies. R. Sayre supplied data on component operating temperatures and section thicknesses. 7 l-4 . 4' e 'g-
*4
=
s7.;
-ji i
Xi
- 1. INTRODUCTION A study has been made of the pr:certies of Types 304 and 316 stainless steel in order to critically assess their relative merits for coolant-boundary applications in liquid metal fast breeder reactor (LMFBR) systems.
. Materials in the low, normal and high carbon grades are considered. The i objective of this evaluation is to make recommendations for the selection of materials for. components in the primary, intermediate and auxiliary I systems of the CRBRP based on material performance and, to a lesser extent.
[ cost. For the steam generator an additional evaluation has been made of the properties of 2-1/4 Cr-1 Mo low alloy steel which will be used in this component. The components of interest include, but are not necess&rily limited to, the following: Primary Syst_em a )' ~ Vessel b) Vessel outlet nozzle c) Vessel inlet nozzle d) Hot-leg piping e)_ Cold-leg piping f) Pump suction nozzle g) Pump discharge rozzle h) Pump casing i) Intermediate heat exchanger inlet nozzle j) Intermediate heat exchanger outlet nozzle k) Intermediate heat exchanger shell
- 1) Hot-leg isolation valve inlet nozzle
" m) Hot-leg isolation valve outlet nozzle n) Valve body 7 -o) Cold-leg isolation valve inlet nozzle -} p) Cold-leg isolation valve outlet nozzle q) Cold-leg isolation valve bcdy 1
s Intermediate System a) Intermediate heat excnanger inlet nozzle b) Intermediate heat exchanger-outlet nozzle c) ' Hot-leg piping d) Cold-leg piping e) Pump suction nozzle f) Pump discharge nozzle g) Pump casing h) Hot-leg isolation valve
- 1) Hot-leg isolation valve outlet nozzle j) Hot-leg isolation valve body k) Cold-leg isolation valve inlet nozzle
- 1) Cold-leg isolation valve outlet nozzle m) Cold-leg isolation valve body n) Steam generator tubing o) Steam generator inlet nozzle p) Steam generator outict nozzle q) Steam generator shell Auxiliary Systen_1 a) Auxiliary system hot-leg piping b) Auxiliary system hot-leg sample loop piping For a particular application it is not always possible to differentiate between the behavior of Type 304 and Type 316 stainless steel. For example, in sensitization, stress-corrosion cracking, fabrication and weldability the differences between the two materials are small. The selection of one over the other is, therefore, often made on a lowest-
- cost basis. More importantly, in this evaluation, significant emphasis is placed on determining the merits of allcys containing various starting contents of carbon and nitrogen since it is vital to estimate the long-term effects of interstitial transfer on mechanical behavior in an LMFBR system. 2
Property evaluations for tne various materials are given in the following sections.
- 2. PROPERTIES OF TYPES 304 AND 316 STAINLESS STEEL _
2.1 Mechanical Properties 2.1.1 Short-Term Tensile Properties The mechanical properties of the unstabilized austenitic stainless steels are influenced by the presence of interstitial elements such as carbon and nitrogen. Although it has been demonstrated that nitrogen has as strong an effect as carbon,II) and most Type 304 and Type 316 stainless steel contain more nitrogen than carbon (Figure 1), the ASME Boiler and Pressure Vessel Code does not address the problem of nitrogen content in stainless steel. Class-ifications within a particular grade of steel are given in terms of carbon content only (see Table 1). For high temperature service (> 800 F) low-carbon grades of material are not permissible and the minimum carbon content for normal grades is specified as 0.04 weight percent.(3) Depending on the carbon content of the material, different allowable design stresses may be specified in the ASME Code. It is, therefore, important to assess whether the interstitial content in a sodium-exposed reactor component falls below specified minimums during service and also to determine the impact of changing interstitial con-tent on mechanical properties. 2.1.1.1 Effect of Carbon and Nitrogen A comprehensive analysis has been, completed on the effects of carbon and nitrogen concentrations on the yield strength, tensile strength and total elongation of solution-treated Types 304 and 316 stainless steel.U) Since most of the available data in the literature do
- not include nitrogen contents, the amount of useful information available is small, especially for Type 316.
3
The following assumptions and guidelines were used in selecting data for analysis in reference 1: a) _ All alloys must have chemical compositions within the limits specified by the American Iron and Steel Institute (AISI). 7
- i Plate, bar, pipe, tube, and forged materials were considered.
b) c) Carbon and nitrogen were assumed to have equal and additive effects on mechanical behavior; hence, the total interstitial content may be given as (C + N). For convenience, C and N' t are given in weight percent although it is more correct to use atom percent. d) Alloy grain size was assumed to be a negliaible variable in the analysis. ' ' e) The test strain rate was assumed to be a negligible variable in the analysis. f) Only solution-treated material was considered. The solution temperatures were between 1900 and 2100 F, with most alloys being water or air quenched. g) In developing the design equations the mathematical form were generated, where possible, from existing mechanical behavior theories. This increased confidence in extrapolating the equations to very low (C + N) levels where no data exist. Where theory could not be utilized empirical data fits were made.
- J 4
9 For Type 304 stainless steel the yield strength, ultimate tensile stren9th and total: elongation cata were shown to be(I) respectively: oy= 89.952 + 181.167 (C + N) - 0.148T (C + N) 1/2 q - 5.727 T + 0.105T (1) a-p standard deviation = 2.236 P [ Valid for 297 < T < 922 K, 75-1200 F] l u
= - 1.806 - 227.391 (C + N) + 1218.392T
-4
+ 1.127T (C + N) - 8.446 x 10_
T2 (C + N) (2) ( standard deviation = 2.645
-[ Valid for 366 < T-< 922 K, 200-1200 F]
-1/2 e = 29.595 - 89.898 (C + N) + 468.414T (3) standard deviation = 2.955
[ Valid for'477 < T < 922 K. 400-1200 F] These equations.are valid to a maximum interstitial concentration of 0.13 weight percent. The yield ar,d tensile strengths are in ksi, the total. elongation in weight percent and the temperature in degrees Kelvin. 5
e i For Type 316 stainless steel the equations are: i (4) c y = 88.474 + 211.640 (C + !{) - 0.206T (C + tt) 1/2
- 5.762T + 0.111T standard deviation = 2.806
-1/2 u
= 8.563 - 461.665 (C + ft) + 1064.154T
-3
+ 2.722T (C + ft) - 2.330 x 10 T2 (C + ft) (5) standard deviation = 6.458
-1/2 c = 28.000 - 99.139 (C + ft) + 653.926T (6) standard deviation = 6.703 The above equations are usually valid for temperatures between 297-977 K (75-1300 F) and (C + ft) 4. 0.010 weight percent. How-ever, at room temperature, 1100 F and 1300 F, the range of validity may be increased, respectively, to interstitial levels of 0.167, 0.141, and 0.141 weight percent.
In Figures 2 through 5 are given direct comparisons between Type 304 and. Type 316 stainless steel for various (C + ft) levels. For equivalent interstitial levels there is very little difference in the yield strengths of the two alloys but the ultimate tensile strength and' ductility of Type 316 stainless steel are markedly higher. However, for the lower interstitial levels the differences in the ultimate tensile strength of the two materials becomes smaller. 6
An important observation in Figures 3 and 4 is that a Type 304
~
or Type 316 steel containing a minimum required 0.04 weight percent carbon alone would not meet the specified minimum room temperature yield strength of 30 ksi. However, alloys containing an additional 0.04 weight percent nitrogen would meet this require-ment. This clearly illustrates the necessity to know the nitrogen p as well as the carbon content in an austenitic stainicss steel. 3 In Table 2 is given the minimum anticipated yield strengths (S y) and the time-independtnt allowable design stresses (S m ) for the various grades of Type 304 and Type 316 stainless steel. For normal and high carbon. grades the yS and Sm values for Type 316 stainless steel are higher. For the low-carbon grades, values for Type 304 are higher. However, if the effects of nitrogen are considered, as was the case in Figures 2 through 5, the yield strength of Type 316 stainless steel is always marginally higher. Hence, if a direct comparison is made between the tensile behavior of Types 304 and 316 stainless steel for low, normal and high interstitial grades " Type 316 is superior. For any given (C + N) concentration it has higher strength than Type 304 and it has approximately 5 percent more ductility over a wide range of temperature (Figures 2 through 5). The above discussion has centered on solution-treated material. Thenn&l aging causes the precipitation of carbides and intermetallic phasesI4} and results in changes in mechanical behavior. Avail-able data indicate that there is a small increase in yield strength a zero or marginal increase in tensile strength and a decrease in total elongation.(5, 6) 2.1.2 Creep Properties The creep properties of Types 304 L, 304, 304 H, 316 L, 316 and 316 H stainless steels have been compared by Smith.(7) Most of the data are for temperatures of 1000 F and higher since creep at lower temperatures is extremely slow. A sumary of the creep properties in the range 1000 to 1100 F is given in Figures 6 and 7 No correlation was made with nitrogen contents but it may be seen 7
that L grades of Ty e 304 and Type 316 have faster secondary creep rates than the regula- carbon grades. Type 316 stainless steel has a slower creep rate than Type 304 at any given stress level but at 1100'F the creep rates of Type 304 and 316 L are similar. The rupture strengths of the L grades are less than those for the > regular grades. Type 316 has a higher rupture strength than Type 304 but Type 304 and Type 316 L behave similarly.
- Rupture strength data which take into account nitrogen, have been plotted for Type 304 and Type 316 solution annealed bar and are given ir Figures 8-11.(8) Although both 103 and 105 hour rupture strength data are plotted the latter were obtained by extrapola-tion and, hence, may be subject to some error. The following 9eneral conclusions may be drawn from the data in these figures
a) Both carbon and nitrogen concentrations must be specified since both elements strongly influence creep behavior. b) Based on the small degree of scatter in the data, to a reasonable approximation the nitrogen and carbon effects are of equal magnitude so that the effective interstitial concentration may be written simply as (C + N). c) Increasing (C + N) content increases the rupture strength. With the exception of the 103 hour rupture strengths for Type 316 at 1100 and 1200 F, the rupture strength is a linear function of the (C + N) content within the concentration limits evaluated. 1 d) For equivalent (C + N) concentrations and test temperatures 4 the 103 hour rupture strengths of Type 304 stainless steel are about 20-30 percent less than those for Type 316. The 1 105 hour rupture strengths of Type 304 are about 40 percent less than those for Type 316. 8
e) The effect cf (C + N) concentration on rupture strength decreases with increasing test temperature. A limited amount of work has been conducted on the uniaxial creep of Type 316 stainless steel which was.decarburized and denitrided in flowing sodium prior to testing.(9) From stress relaxation tests at 1325"F for stresses below 13 ksi the treep rates were obtained and are given in Figure 12. For the sodium-exposed material tnere 1 was a loss in carbon and nitrogen from 0.0558 and 0.0556 weight
- percent, respectively, to 0.0466 to 0.0106 weight percent, res-pectively. Compared with the argon-aged material a (C + N) change from 0.1114 to 0.0572 increased the creep rate by an order of magnitude.
All of the above data clearly show the large effect which carbon and nitrogen transfer will have on the creep behavior of unstabilized austenitic stainless steel in an LMFBR environment. 2.1.3 Fatigue Properties A comparison between the fatigue behavior of Type 304 and Type 316 stainless steel has been made for temperatures between 806 and 1202 F. 00) The Type 304 contained 0.053 and 0.052 weight percent carbon and nitrogen, respectively, and the Type 316 0.06 and 0.048 weight percent carbon and nitrogen, respectively. For equivalent heat treatments and test temperatures Type 316 has a shorter If the fatigue life than Type 304 at the same total strain range. comparison is based on the stress amplitude, however, Type 336 is superior. This is because for a given stress the corresponding strain range for Type 316 is somewhat smaller than that for Type 304 owing to the higher yield strength of Type 316 stainless steel. Hence, a selection of Type 304 or Type 316 for fatigue resistance will depend on whether the fatigue condition will be strain or 1
- stress controlled.
9 w w v .- ,-,y--- % ,.
-r- wr %- --- - .,.-. -
--r--- - - - - - -,--,--w -sv----,, ,--.---,w- - ----w4r
V O The effect of carbor content on the fatigue behavior of 18 Cr-12 Ili
}
iron-based alloys his Deen studied by Driver. It was shown that at elevated te :eratures in the range 1112 to 1472 F, for a wide range of stress amolitudes, decreasing the carbon level from 0.05 to 0.004 weight percent causes a much shorter fatigue life and endurance limit. This was attributed to carbides effectively blocking grain-boundary sliding and migration, thereby inhibiting the nucleation of grai, boundary cracks. At room temperature, however, it has been shown that for Type 316 stainless steel fatigued at strain amplitudes in the range + 1 to + 4 percent the presence of 2 to 3 volume percent of chromium carbide can decrease the fatigue life to about one-third of that in solution treated material.02) This is caused by the brittle fracture of the grain-boundary carbides. There are no data for comparisons in the
- current temperature range of interest, viz. 650 to 1000 F, hence the effect of interstitial concentration is not known with certainty. . 2.2 Sensitization And Heat Treatment Effects An unstabilized austenitic steinless steel is said to be sensitized when carbides are precipitated at grain boundaries during high-tem-perature exposure. Precipitation of the carbides occurs readily in the temperature range 800-1600 F.03) / \
When the degree of sensitization is such that continuous carbide paths are formed at the grain boundaries intergranular corrosion may result if the material is exposed to a corrosive atmosphere. A simple theory to explain this phenomenon is based on the formation of continuous Cr C layers at the boundaries which results the formation of 23 6 In these regions adjacent zones which are denuded of chromium. therefore, a decreased resistance to corrosion occurs and inter-granular attack may occur. Particular agents which have been reported to cause inte granular attack in sensitized austenitic stainless steels are given in Table 3. In the early stages of carbide precipi-tation, however, when continuous grain-boundary carbide paths have not been established, significant intergranular corrosion is not likely. 10
1 2.2.1 Measurement of Sensitization The degree of carbide precipitation is dependent on both the time and temperature of exposure. At low temperatures (s 800*F) precipitation is slow because of the slow rate of diffusion of
, chromium to the grain-boundary carbides. At high temperatures (s 1600*F) the de9ece of supersaturation of carbon in the alloy is relatively low and there is less tendency for carbides to be l t
formed. The tests usually employed to detect sensitization are described in ASTM procedure A262.(15) Typical sensitization curves obtained are given in Figure 13. The data represent the corrosion rates in boiling Strauss solution (acidified copper sulfate) for material pre-exposed at the temperatures and times shown. flote that for exposure temperatures of about 1300 F sensitization, manifested by increased corrosion rates, occurs at shorter exposure times. It is also clet.r that for the low-carbon material sensitization is significantly delayed. 2.2.2 Occurrence of Sensitization in LMFBR Components Sensitization is likely to occur during fabrication, stress relievii.g and service of most of the components listed in Section 1 of this report. During fabrication, for example, the cooling of components after hot forming and welding will result in passing through the sensitizing temperature range. For heavy section comprents it may not be possible to quench rapidly enough, so sensitization is inevitable. Stress relieving is necessary in many structures to ensure dimensional stability during service. The data given in Table 4
- show the degree of stress relief which can be achieved in N austenitic stainless steels for various annealing temperatures.
Ideally, stress relief should not be conducted in the 800 to 1600'F sensitization range. If possible, it should also be
. avoided at high temperatures since necessary rapid cooling through the sensitization range might restore residual stresses trd cause distortion.
11
O O In the case of sensitization during plant operation many of the components listed in Section 1 of this report are designed to operate over a 250,033 hour lifetime at temperature up to about 1050*F. Under such conditions sensitization is unavoidable, as Figure 13 shows. 2.2.3 Effects of Sensitization I The presence of a sensitized structure is itself not a serious ) problem. Compared with c solution-treated material sensitiza-tion in an alloy will cause only small increases in strength and small ductility losses.(5. 6) However, as shown in Table 3 a particularly large range of corrosive agents will cause some degree of grain boundary attack in a sensitized alloy. Rusting, for example, is quite common in the heat-affected zones of welded components. Probably the most important effect of sensitization is that during fabrication, shipping or erection of LMFBR components intergranular corrosion will form incipient cracks which may act as stress raisers. This could induce stress-corrosion cracking for certain environments which the cc:aponent is later exposed to. This problem is discussed in Section 2.3. 2.2.4 Comparison of Alloys for Resistance to Sensitization Effects A straightforward comparison between Type 304 and Type 316 for resistance to sensitization-induced corrosion is difficult for the following reasons: a) Data for alloys containing the same carbon concentration are scarce, b) There is difficulty in obtaining data on alloys containing typical concentrations of alloying elements, especially chromium. 12
c) The presers of nitrogen could influence sensitizatio ects. Work reported by Hazelton indicates that nitrogen ons resulted in decreased amounts of sensitization. Data which support this conclusion are given in Tables 5 . From fundamental considerations the higher chromi Y# Type 304 would tend to reduce the extent um contents in of chromium zones at the grain boundaries. epleted 316 has % Also, molybdenum additions (Type
?. 5 percent molybdenum) encourages sensitizatio at temperatures below 1200 F(23) (Figure 14) .
to support the contention that Type 304 stainlessExperim steel is less
'
- susceptible to sensitization is given in Figures or 15 and low-carbon materials.
grades of steel, hence it may be ype a prematur 304 is definitely superior under all possible sensitizi ng conditions. Under certain situations it may be possible to c onsider weaker low-carbon grades of steel to overcome sensitization problems. i, This may be feasible, for to the highest stresses or where operating tempe Figure 13 shows the beneficial effects of ratures low care low. delaying the onset of sensitization. - arbon material in 316 L have been heated at 1200 F for 1 to 3 days a ave shown no intergranular attack in the ASTM-262 test after th ree successive 72-hour immersions in the StraussHence, . solution the use ( 5) of low-carbon materials are of significant benefit sinc e there is far greater latitude in welding and stress relievin conducted at high temperatures. g operations For normal and high-carbon grades stess relief below abcut 950*F may be required b will only reduce peak stresses b u this 2 to 4 hour annealing perica.(I about 25 to 30 percent in a treatment is mandatory an anneal above 1600 F isIf a required (see Table 4) followed by as rapid a cool as possible rough the th 13
sensitization range. The permissible cooling rate will depend on the section thickness since care must be exercised to avoid re-introducing residual stresses and distortion. 2.3 Stress-Corrosion Cracking Stress-corrosion cracking is a recurring problem for aust.enitic stain-less steel components in the nuclear industry. It is a brittle failure phenomenon caused by the conjoint action of stress, environment and possibl'e temperature. In references 26 and 27 a review of case histories ] and laboratory tests showed that the following agents were possible contributing causes of stress-corrosion: chlorides, fluorides, caustic, i tar oils, natural oils, organic solvents, orange ju1ce, milk, welding flux, amonia, sulphates, acetic acid, alcohols, corn syrup, sulphides,
; baked beans, coffee, tomato soup and body fluids, including blood serum.
j Of these the greatest proportion were attributed to chloride-induced failures and most of the remainder were from fluoride or caustic contam-S ination.(26) In the LMFBR industry all three contaminants may be encount-j ered at some stage of fabrication, shipment, erection and service operations. 2.3.1 Effect of Environments on Chloride Stress-Corrosion Cracking A large amount of work has been conducted on chloride-induced stress-corrosion and a review is given in reference 13. It appears that moisture is an essential requirement for cracking.03) A standard test for evaluating resistance to chloride-stress corrosion is to expose stressed samples to boiling 42 percent magnesium chloride solution. This is an extremely aggressive environment and it bears little similarity to conditions which might be expected in industry. However, it serves as a useful ]- indication of the relative behavior of alloys which are sus-l ceptible to chloride cracking. Edeleanu(28) has shown that in boiling aqueous solutions cracking is accelerated by increasing magnesium chloride 14
concentration (Figure 17). Other investigations have demon-strated that the presence of oxygen accelerates chloride-induced failure.03) 2.3.2 Effect of Temperature on Chloride Stress-Corrosion Cracking Several studies have shown that increasing temperature accelerates failures in chloride environments.0 3) The time to failure (t ) fappears to be related to temperature by an / e,wression of the form: t f= A exp (B/T) in which T is the absolute temperature and A and B are positive constants. ) 2.3.3 Effect of pH Value on Chlottde Stress-Corrosion Cracking All of the available data indicate that increasing the pH value l leads to increased time to failure in a given chloride environ-ment under a particular set of conditions for applied stress, chloride concentration and temperature.0 3) 2.3.4 Effect of Alloy Content on Chloride Stress-Corrosion Work by Copson(29) indicates that for increasing concentrations above approximately 10 percent, nickel leads to enhanced resis-tance of iron-chromium-nickel alloys to stress corrosion (see Figure 18). In the permissible range of nickel concentrations for Type 304 and Type 316 stainless steel (8-10 and 10-14 percent, respectively), the data are not explicit but it does appear that Type 316 stainless steel should resist stress-corrosion more { effectively. The data in Figure 19 clearly confirms this. These same data indicate that for regular and low-carbon grades of { I Types 304 and 316 stainless steel the resistance to cracking is 15
e. similar. However, there is some indication that for very high carbon concentrations, above about 0.1 weight percent, increased resistance to stress corrosion is likely (see Figure 20). Nitrogen additions appear to have the reverse effect and enhanced susceptibility to stress corrosion has been reported.(32, 33)
. 2.3.5 Effect of Structure on Chloride Stress Corrosion Cracking The effects of sensitization on cracking have been investigated but the data are conflicting. .Scharfstein and Brindley(34) found that for a given exposure in a chloride solution sensitized Type 304 showed deep intergranular cracks. As-received material showed shallower transgranular cracks. Rideout(35) confirmed the deleterious effects of sensitization. Other workers, how-ever, find that for the austenitic stainless steels the degree of sensitization does not noticeably affect chloride cracking,(30,36,37) with the exception of Type 304 L which appeared' to be adversely affected by sensitization.(37) 2.3.6 Stress Corrosion in Caustic Environments Compared to chloride-induced cracking relatively little work has been conducted on caustic stress corrosion. This type of failure is important in an LMFBR system where caustic contamination may occur in the intermediate sodium system because of leakage in the steam generator. Recent work (38) on Types 304 and 304 L stainless steel has studied the nature of. cracking in 10 and 50 percent sodium hydroxide solutions at 600 F. For "U-bend" samples made from material in the as-received and the solution-treateci conditions both the 304 and 304 L grades failed in less than 3 days in the 10 percent solution and in less than one day in the 50 percent solution.
Under the inspection procedures used, no differences could be detected between the two grades of steel. In the 10 percent solution cracking was intergranular but for the 50 percent solution it was transgranular. 16
Aging of the as-received and solution-treated materials for periods greater than about 100 hours in the sensitization range were effective in inhibiting cracking in the 10 percent soietion but had no apparent effect in the 50 percent solution. Work by Agrawal and Staehle(39) indicates, however, that in boiling solutions of sodium hydroxide at ambient pressures sensitization is detrimental in Type 304. Reasons for the discrepancy are as yet unresolved. 2.4 Sodium Corrosion Rates In the high-temperature regions of flowing sodium systems austenitic stoinless steels are susciptible to the loss of metallic elements such as chromium, nickal and manganese.(13) The loss of nickel and manganese, which stabilize austenite, results in the formation of a ferritic surface layer which will be substantially weaker than the bulk material. As the ferritic layer gets thicker the rate at which the metallic elements diffuse through it becomes slower and a steady state is reached when the proportions of elements removed by the sodium is approximately equal to the chemical composition of the si. eel . Estimates have been made for the steady-state ferrite layer thickness for an all-austenitic system at a maximum tempera-ture of 1100 F.(40) For an oxygen content of < 10 ppm, as measured by the amalgamation technique, and a sodium flow rate of 7 to 20 feet per second, the thickness is about 0.13 mils. This will not cause a measurable effect on the overall strength of an LMFBR com-ponent. Estimates have also been made for the corrosion rates of austenitic stainless steels. For conditions similar to those given above a steady-state rate of 0.08 mils / year was obtained.(40) This translates to a net corrosion loss of 2.4 mils over a 30-year life-time for a component at 1100 F. For the components listed in Section 1 the corrosion losses would be smaller because of their lower operating temperatures. 17
2.5 Fabrication and Weldability A detailed review of fabricanility and weldability of Type 304 and Type 316 stainless steel is given in reference 13. Differences in behavior between the two are relatively minor and a selection 7 of one material over the other based on these properties would be difficult. I ^ Differences do exist, however, between the high, low and normal carbon grades. As stated in Section 2.2, sensitization problems are greatly minimized for low-carbon alloys which allows greater latitude in stress relieving heat treatments which may be required. Welding problems ane also decreased for the same reason. The order of preference, for minimizing welding difficulties, is 304 L, 304, 316 L and 316.0 3)
- 3. PROPERTIES OF FERRITIC 21/4 Cr-lMo STEEL 2-1/4 Cr-1 tio steel is used at temperatures up to 1200 F in situations where high corrosion resistance is not required. Table 7 gives the chemical composion limits for this material. The chromium content improves corrosion re-sistance and the molylodenum increases the yield, tensile and creep strengths.
If this alloy is rapidly cooled from hot forming or welding temperatures a hard brittle phase called martensite is formed which greatly decreases ductility. Heat treatment is required to restore some of the ductility. Two basic annealing . treatments may be used.(4I)- a) Soak material at 1500 to 1700 F and slow cool. b) Soak material at 1200 to 1400 F and air cool. e The length of time required for soaking increases with increasing section thickness. 18
In order to use the material in a higher strength condition it may be either normalized and tempered or quenched and tempered. For the first case normalization involves annealing in the range 1650 to 1750 F followed by air cooling. Tempering is achieved by reheating to 1200 to 1400 F followed by air cooling.(4I) j For the second heat treatment water-dip-quenching from about 1750*F is followed by tempering in the range 1000 to 1200 F followed by air cooling (42) 1
, Properties of 21/4 Cr-lMo are given in the following sections:
3.1 Mechanical Properties. The mechanical properties of 2-1/4 Cr-1 Mo are reasonable well documented but, because of the wide range of practical heat treatments which may be used, substantial variations in properties are noticed. Carbon levels are usually close to the permissible maximum of 0.15 weight per cent because of the enhanced strength. A recent publication, however, has investigated the effects of low carbon levels on the mechanical behavior (43) A. limited amount of work has also been fonned on the effects of in-sodium interstitial transfer on properties.(5, 6) Effects of variations in nitrogen content have not be investigated. 3.1.1 Short-Term Tensile Properties Figure 21 and 22 show typical tensile properties for material in the annealed, nonnalized and tempered, and quenched and tempered conditions. Note the sharp drop in strength above 1000 F and the ductility minimum at 800 F which are typical of all three heat treatments. Information on thennal aging shows that losses in strength will occur although ductility re-mains essentially constant (5, W , In Figure 23 the effects of carbon on yield and tensile strength are given ) . For a heat of material cast with a carbon content of 0.020 weight percent the strength is similar to that for standard material. However, for material decarburized to 0.021 weight 19
per cent a large strength loss is noticed. The difference in behavior-for the te heats is believed to be connected with grain size. Information on in-sodium decarburization shows that, for normalized j- and tempered material, large strengtn decreases and ductility increases are likely.(5) 3.1.2 Stress-Rupture Properties 4 Figure 24 compares the 10 3
,10 and 105 hour rupture strengths for annealed and normalized and tempered 2-1/4 Cr-1 Mo.(74) The normalized and tempered material is significantly superior to !
annealed material. Some information on the 1100 F stress l rupture properties of normalized and tempered alloys, as a function of carbon, is summarized in Figure 25. For a given stress, decreasing the carbon level from 0.13 to 0.027 weight percent will decrease the rupture time by an order of magnitude. This is equivalent to a reduction in rupture strength of about 40 percent (see Figure 26). 3.1.3 Fatigue Properties In Table 8 are shown the low-cycle fatigue properties of standard and decarburized mattrials at 900 and 1100 F for a total strain ranga of 0.005 in/in at a frequency of 0.25 Hz. For low-carbon concentrations a drastic decrease in fatigue life may be anticipated. 3.2 Stress Corrosion Cracking b There do not appear to be any published data which clearly show that y unstabilized 2-1/4 Cr-1 Mo steel is susceptible to stress-corrosion cracking in chloride or caustic environments. 3.3 Sodium Corrosion Rates The corrosion rates of 2-1/4 Cr-1 Mo have been measured by several groups over a wide range of temperatures and oxygen levels. Figure 27 summarizes the available information. None of the oxygen analyses 20
O were apparently measured by the currently favored vanadium wire technique. However, with the exception of the 0.5 ppm oxygen analysis, which was obtained by the plugging meter technique, the 1100 F corrosion rates for 21/4 Cr-1 Mo appear to be higher than those for Types 304 and 316 stainless steel (see Section 2.4). The work of Sannier et al(75) ) directly compared the corrosion rates of 21/4 Cr-1 Mo and Type 304 stainless steel for exposure periods of about 6000 hours at 1200 F. The
' data, given in Figure 28, show that 2 1/4 Cr-1 Mo has a much faster initial corrosion rate although at longer times the steady state rates for the two groups of alloys may not be significantly different. Hence, for LMFBR coolant-boundary components the metal corrosion loss at maximum operating temperatures is not likely to be a problem over the 30-year lifetime of the components.
3.4 Steam-Corrosion Rates The data presented in Figures 29 and 30 show the steam corrosion rates for 2 1/4 Cr-1 Mo for temperatures between 950 and 1200 F. Specimen conditions, steam pressures and boiler water chemistries are given in Table 9. All exposures were under isothermal conditions. The gradient of the 950 F line was predicted by assuming that an Arrhenius type oxidation rate was applicable, thus enabling an extrapolation of the oxidation rates for the 1100 and 1000 F tests. Conclusions obtained from these data include: a) Within normal experimental scatter there is no difference between the corrosion rate of as-machined, pickled, and pickled and stress-relieved surfaces. . b) Within the range of steam pressures studied the corrosion rates appear to be generally independent of pressure. c) There is a large difference in the corrosion rates between Oak Ridge National Laboratory and ASME samples at both 1100 and 1200*F. One possible explanation is that the ORNL tests 21
O involved reducing :ne specimen temperatures to room temperature every 1000 hours in order to expose them intermittently to saturated steam. Inis could possibly cause cracking of the protective oxide scale which would lead to accelerated Corrosion. d) At 1200 F the rate of corrosion is markedly accelerated. t Table 10 gives estimates of the metal loss for 30-year exposures of 2-1/4 Cr-1 Mo to steam. 3.5 Sodium-Water Reaction Effects An early study compared the rates of wastage for 2-1/4 Cr-1 Mo, Incoloy 800 and Type 316 stainless steel during exposure to a sodium / water reaction.(51) This type of information is most important in selecting LMFBR steam generator tubing material because of the possibility of a water-to-sodium or steam-to-sodium leak. Table 11 gives the wastage rates for these tests. The testing technique involved the injection of water into a pot containing a 3 lb. mass of sodium at 930 F. The water was injected at right-angles to a plate sample which was weighed before and after a 30 minute test. The 2-1/4 Cr-1 Mo was greatly inferior in performance to the austenitic alloys. More recent studies i have confirmed the poor resistance of 2-1/4 Cr-1 Mo to caustic corrosion.(52) 3.6 Fabrication and Weldability 3.6.1 Hot Forming Suggested temperatures for bending are 1750 to 1850 F and for forging, 2000 to 2200 F. Stress relieving should be conducted at 1300 to 1400 F and annealing should be at about 1000 F followed by slow cooling. Some ferritic steels, including 2-1/4 Cr-1 Mo, may, under certain conditions, be susceptible to temper embrittlement and stress-relief cracking. These phenomenon are, as yet, incompletely 22
understood but procedures have been developed to minimize their occarrence during fabrication and sub-sequent component service. A detailed review of these and associated embrittlement problems for 21/4 Cr-1 Mo is given in reference 53. 9 3.6.2 Weldina(54)
- 0 2 1/4 Cr-1 Mo is readily weldable by the shielded metal-arc, submerged-arc, gas metal-arc, flux-cored-arc, gas tungsten-arc and electroslag process
- s. Since this at 'oy is air-hardenable, care must be exercised to prevent cracking I of the weld and heat-affected zone materials. Minimum pre-heat temperatures of 300 F are usually specified to reduce the possibility of cracking by:
a) Allowing hydrogen to escape, if it is present. I b) Reducing the rate of cooling from the welding temperature. Filler metals containing 2 1/4 Cr-1 Mo may be used although United States Steel Corporation have recommended 5 Cr-1/2 Mo.(4I) Post-weld heating to 1100 to 1375 F is usually recommended for stress relief. 3.6.3 Welding of 21/4 Cr-1 Mo to Austenitic Stainless Steels The joining of ferritic to austenitic materials poses special problems because of the different thermal expansion rates; see Table 12. In addition, carbon will tend to diffuse from the higher-carbon ferritic material during service into the weld metal which will have a composition dependent on the proportions 23
- q b of melted ferritic, austenitic and filler metals. Hence, problems may arise from the higher carbon content which will occur in the austenitic material, or the depleted carbon level in the ferritic steel, after long-tenn service at high temperature. Such weld failures have continued to recur even after 20 years' of experience.(57-59) Filler metals which have been satisfactorily used are: Type 307,16-18-2, Inco 182(60-63) > (i.e. ENiCrFe-3). Chemical compositions of these alloys are given in Table 13. To ensure dimensional stability, dissimilar metal joints should be heated to 100 F above the maximum operating temperature and slow cooled.(13) A detailed welding procedure for welding Type 304 stainless steel pipe to 2 1/4 Cr-1 Mo has recently been described.(64) Because of its direct relevance to the current project, the description is reproduced below in its entirety: Nickel Alloy Filler _ Metal For__ Crack-Free Welding The selection of a filler metal was an important factor in welding a 2 1/4 Cr-1 Mo steel pipe (ASTM A387, grade D) to a Type 304 stainless steel pipe. This pipe assembly was a com-ponent of a steam pipeline that had to withstand cyclic heating and cooling between 1050'F and room temperature. Details of the joint between the 15-in.- OD by 1-in. wall pipes before and after welding, in 22 passes, by the shielded metal-arc process are shown in Figure 31. A nickel alloy flux-covered electrode (ENiCr Fe-3) was selected in pre-ference to an austenitic stain-less steel electrode, for three reasons: 24
a) The coefficient of ther al expansion (8.5 nicro-in. per inch per *F) of weld metal from the ENiCr Fe-3 electrode is close to that of the ferritic 21/4 Cr-1 Mo steel. Thus, during cyclic temperature service, the major differential expansion stresses developed primarily at the stronger interface between the stainless steel pipe.and the weld metal, rather than at the weaker interface between the ferritic steel pipe and the weld metal - the location at which such expansion stresses would have developed if a stainless steel elec-trode had been used. b) Carbon depletion in the ferritic steel was less when the nickel alloy weld metal was used than when a stainless steel weld metal was used. Therefore, the heat-affected zone of the ferritic steel was not weakened by loss of carbon. 1 c) The excellent depo-sition characteristics of the nickel alloy 9 electrode produced a sound, porosity-free weld that was not subject to weld-metal cracking. The pipe ends were machined to the joint configuration shown in Figure-31 and wiped clean with a solvent; then_the pipes 25
were mounted horizontally on turning rolls for welding. Low-frequency induction heat-ing was used to provide a 500 F preheat for the ferritic pipe. This temperature was maintained, by automatic control, throughout the welding operation. The. joint was-tack welded ! I at 6-in. intervals, using a 1/8-in diameter elec-trade and the same current and voltage settings as for the first pass (see table of welding conditions with Figure 31. The edges of the tack welds were thinned by grinding before the root pass, which was carefully deposited to obtain a smooth inside root surface. Each weld bead was deslagged and visually inspected before the next bead was deposited. Immediately after welding, the weld was postheated, by induction heating, to 1350 F for 1 hour and air cooled. Welds made with the ENiCr Fe-3 electrode met bend-test and tension-test requirements of Section IX of the ASME Boiler and Pressure Vessel Code, and also passed visual, dye-penetrant, radiographic and metallographic inspec-
, tion. Fractures produced in transverse tension tests at room and elevated tem-peratures were in the base metal away from the heat-affected zones. The welds had markedly better resis-tance to cracking during thermal cycling than did similar joints welded with
-Type 347 stainless steel electrodes.
26
- 4. INTERSTITIAL TRANSFER EFFECTS The transfer of interstitials, sz h as carbon and nitrogen, in LMFBR components is a potential problem because of the attendant changes in mechanical strength during long-term service (see Sections 2.1 and 3.1, above). Conceivably, a component meeting minimum strength and interstitial content requirements at start-of-life could violate these requirements at some point during service.
To date, there are no proven techniques which will accurately predict the rates of interstitial transfer and end-of-life mechanical behavior under complex LMFBR conditions. However, by using various simplifying assumptions attempts have been made at ARD to derive estimates of the extent of interstitial transfer.(66-68) 4.1 Selection of Carbon Potential Values In order to estimate the rates of interstitial transfer in the CRBRP primary and intermediate systems, values for the carbon potential (Cs) need to be established. ARD is currently assuming that the carbon potential for both systems is 30 ppm. Reasons for this selection are given below: a) Primary Side AP.D had carried out a survey of the carbon potentials of operating test loops fabricated of unstabilized 300 series stainless steels, including EBR-II.(69) Generally, the survey indicated that the carbon potentials were low and of the order of 30 ppm. The 30 ppm Cs value has been used for FFTF pre-dictions (66) and is likewise considered to be the most repre-3 sentative value for estimation purposes here.
)
b) Intermediate Side Early work at G.E.(70) and some observations made by B & W(71) indicate that when a system contains a ferritic alloy, such 27 t-
as 2-1/4 Cr-1 Mo, tne carbon lost from the ferritic alloy will cause carburization of the stainless steel in the system. Recent G.E. work, however, would suggest that decarburization rates of 2-1/4 Cr-1 Mo are much lower than previously expected (72) . These observations are supported by measurements made at ARD on the diffusion of ' carbon from 2-1/4 Cr-1 Mo(69) and by French workers who saw virtually no decarburization below 900*F after 6000 hours (43) However, . no previous estimates of C shave been m;de for bimetallic (austenitic/ferritic) systems. Until direct measurements are made, therefore, it will be assumed that for the inter-mediate system, C s is equal to 30 ppm. 4.2 Compensation for Nitrogen Transfer It is known that nitrogen has at least as great an effect on the mechanical properties of austenitic steels as carbon (I) . It is also known that nitrogen will transfer in a manner similar to that of carbon. The following assumptions are made to evaluate the additional effects of nitrogen transfer: a) The nitrogen content is assumed to be equal to that for carbon. b) The Cs versus Ce relationships for carbon (Ceis the carbon content of the stainless steel at the sodium interface) are assumed to be applicable to nitrogen. These assumptions imply that in all subsequent calculations the effective interstitial content (C + N) is taken to be 2C, and that nitrogen is assumed to behave indentically to carbon. 4.3 Calculation of Interstitial Gradients In order to obtain interstitial concentration gradients in components exposed to sodium for a design lifetime of 250,000 hours (s 30 years) a computer program was used to solve Fick's second law of diffusion (66) , 28
-*--a-7 -ew-. 9,.%7 -
y...__#, __,_ _g. -,~,4-w,-y,,.,_.,_.-,_4.-, ,,,w -- -
,,e.---,4m_ _ -- .
s This states:
'b c = D '2 c 3X E in which c = interstitial concentration, i.e. (C + N) t = diffusion time x = depth below sodium / metal interface F D = effecthe interstitial diffusion coefficient This equation was solved using the boundary conditions x = 0 at c=C e and x -- a a t Lc_ = 0 3x Values of C and D were obtained from reference 66. Component tem-e peratures used in the calculation are given in Table 14 together with materials which were studied in the computer analysis. Typical values of interstitial content have been assumed, as shown. Results of the study are given below.
4.3.1 Primary System Hot-Leg Components Hot-leg components at 995 F undergo severe decarburization and ' denitridation at the surfaces (Figure 32). At the sodium / metal interface the total interstitial concentration is about 160 ppm for Types 316 and 316 H and about 230 ppm for Types 304 and 304 H. These are below the minimum permissible carbon concentration of 0.04 weight percent required for high-temperature design (3) , Significant differences exist between Type 304 and Type 316. Because of the slower interstitial diffusion rates for Type 316(3) the depth of interstitial depletion is only about 60 percent of 29 l
that for Type 304. Ne end-of-life depths of interstitial depletion, as a percentage of component thickness, are given in Table 14 in order to illustrate the severity of interstitial transfer for components of varying thicknesses. The results discussed in this section also apply to the auxiliary system hot leg because of the similar operating conditions. 4.3.2 Primary System Cold-Leg Components The data in Figure 33 show that for Types 304 and 304 L a heavily carburized and nitrided surface will result from a 30-year exposure to sodium. The surface (C + N) concentration for these alloys is about 1460 ppm. Since the carbon concentrations are one-half of these values it can be seen that the surface material of Type 304 L exceeds the permissible carbon concentration maximum of 0.03 weight percent I3) . Normal Type 304, however, falls within the 0.08 weight percent limit (3) Car- . burization depths are given in Table 14. 4.3.3 Intermediate System Hot-Leg Components The data for the hot-leg components in the intermediate system (Figure 34) are analogous to those for the primary system. Types 316 and 316 L steel are superior to the equivalent Type 304 grades in resisting interstitial loss. Compared with the primary system, the depth of interstitial depletion is much smaller, s 4.3.4 Intermediate System Cold-Leg Austenitic Components The data given in Figure 35 and Table 14 for austenitic components are similar to those for the primary system. Again heavy surface - carburization and nitridation are likely although the depth of ) . diffusion is small. For both the primary and intermediate system cold-leg components, the heavily carburized and nitrided surface layer will be far less ductile and under fatigue or bending loads, surface cracking may be induced.II2) 30
S 4.3.5 Intermediate System Ferritic Components The rates of irterstitial transfer for 2-1/4 Cr-currently be predicted using the same type of compute as for the austenitic, stainless steels. From Section 4.1 it seems that any transfer will not significantly affect steam generator components except possible, the tubing . ar- Conclusions reached from ARD studies are(0'): a) at much the same rate asstainless for Type 30aThe steel, b) At temperatures below 1100 F, carbon depletion to fall below an equilibrium level of 500 ppm re of the prevailing carbon activity in the sodium.gardless c)
, At temperatures below 900 F, data extrapolation ind that relativel with the material will have a tendency to carburize n eve f Carburization,y low carbon activity levels in the sodium.
however, will not exceed 2000 ppm. w More recent work, however, indicates that Conclusion not be valid. may (c It is now thought likel burization may not, in fact, occur. 96 that below 900 F car-4.4 Comparisen Between Low,fiormal and High-Carbon Grades For all austenitic components situated in the ghottest re i ons of the primary and intermediate system hot legs, significant losses of inter-stitials will occur from the sodium-exposed surfaces. However, the bulk carbon concentrations will remain abovepercent . the 0 04 weigh minimum (3) if the starting (C + tt) levelsecified in this study (Table 14). are simila and denitridation much more effectively . than Type 3 Hence, if max-imum mechanical strength is required, Type 316 choice. would b e the optimum The presence of the low-interstitial surface layer should 2 present no problems from a mechanical standpoint because of its enhanced ductility. 31
O 1 In cold-leg regions Type 316 also resists carburization and nitridation more effectively than Type 304. At these lower temperatures, however. mechanical strength is a less serious design problem. Thus, the selection of either Type 304 or Type 316 would be appropriate. For maximum resistance to carburization and nitridation, Type 316 would be
! the more suitable of the two. Note that for the lower temperature regions of the cold-leg, where large interstitial gradients are present, it is possible to minimize the gradients by selecting a higher carbon grade of steel. In fact, since the C value e is dependent only on temperature for a given sC , there will be zero interstitial gradient at a location in the system where C, is equal to the alloy interstitial level (C + N),
- 5. COST CONSIDERATIONS A detailed evaluation of the relative costs of fabricating Type 304 versus Type 316 components cannot currently be made. Although estimates were obtained on the costs of starting materials, this would only be a fraction of the component cost because of additional fabrication, welding, inspection and testing. An estimate of these costs is beyond the scope of this study.
To assess the differential between Types 304 and 316 starting alloys, estimates for plate material were considered p3) However, the differential was essentially the same for all product forms. Table 15 gives the costs of Types 304, 304 L, 316 and 316 L alloys. For quantities over 10,000 pounds the cost per pound is decreased slightly. No prices were available for high-carbon grades of steel since they would have to be made to order. Material purchased to RDT Standards will increase the base price by 25 percent whereas material meeting ASME requirements for Section III Class 1 components will increase the base price by 10 percent. Any ultrasonic testing require-ments will add an additional increment of 10 percent on the base price. 32
Compared to austenitic stainless steels, 2-1/4 Cr-1 fio starting material is less expensive. However, the greater degree of difficulty in welding, fab-ricability and heat treatment may rake a ferritic component almost as costly as a stainless steel one. 6.> GUIDELINES FOR ALLOY SELECTION The following guidelines may be used in selecting a particular alloy for a specific LIIFBR application. Table 16 sumarizes these guidelines. 6.1 Selection for Interstitial Transfer and liechanical Behavior In Sectian 4, above, it was shown that in hot-leg components made fron~ regular and high-carbon materials the carbon concentrations at the sodium-exposed surfaces could be reduced below ASf1E Code-specified minimums during service, although bulk carbon concentrations in heavy-section materials would not violate these mininum values (2, 3) Note
-that the equilibrium interstitial concentr.cion at the sodium-exposed surface of a component is dependent only on temperature for a given C 3
value (see Section 4.4) so that it is not possible to avoid low surface carbon levels by choosing a high-carbon steel. In addition, the effects of nitrogen, which exert as large an effect as carbon on mechanical strength (I), are not factored into the Code. It is recommended that in ordering any austenitic stainless steel for use in the CRBRP, the nitrogen levels be at least as high as those for carbon and that interstitial con-tents be in the higher range of permissible concentrations. The Type 316 and 316 H grades of steel are superior to Type 304 and 304 H, especially in rupture strength and tensile ductility. If maximum strength is an overriding design criterion, Type 316 H would be the optimum material to choose since it would limit the ninimum permissible carbon level to 0.04 weight percent (see Table 1). In.the cold-leg, components could reeably be made from normal grades of either Type 316 or Type 304 O nc. at these lower temperatures mech-anical strength is a less F.e .< p f 3 sign requirement. To differentiate between Type ha ano .jpe 316 a trade-off study would be required to compare the lower cost, better weldability and fab-33
ricability, and increased resistance to sensitization of Typa 304 with the superior strength and resistance to stress-corrosion of Type 316. Below about 850 F the S mt allowable design stresses for 2-1/4 Cr-1 Mo are larger than those for the Types 304 and 316 stainless steel. At lower temperatures, therefore, if there is a cost incentive, this alloy may possible be used instead of stainless steel. 6.2 Selection for Fabricability, tieldability and Sensitization For components in lower-temperature regions where the highest-strength grades of steel are not mandatory, significant benefits will result in the selection of regular, rather than high-carbon, grades of Type 304 or Type 316. llelding problems are minimized, and any required stress relieving operations are simplified since sensitization is significantly delayed in lower carbon materials. The weldability of 2-1/4 Cr-1 Mo is inferior to that of austenitic stainless steel in view of the requirement for pre- and post-weld heating. However, this material does not suffer from sensitization problems which are common for the stainless steels. Regarding the potential difficulties of dissimilar metal welds in the Demonstration Plant intermediate system, it would be advisable to minimize the number to be used in order to limit potential problems due to different thermal expansion rates for austenitic and ferritic materials, local interstitial transfer, the possibility of sensitizing the austenitic material during post-weld heat treatment of the weld, and weld repairability. 6.3 Selection for Stress-Corrosion Cracking Resistance Type 316 stainless steel appears to be more resistant to chloriue-induced ( l cracking than Type 304, based on the data in Figure 19. However, the problem is complicated by so many variables that it may not be superior under all LMFBR conditions. Hence, the apparent inferiority of Type 304 should not be used as an argument for disqualifying its use in a component. It is not passible, at this time, to identify the stainless steel for best resistance to caustic-induced stress corrosion. 34 I
g Compared to the austenitic stainless steels, 2-1/4 Cr-1 fio is sub-stantially more resistant te nalide and caustic stress corrosion. 6.4 Tentative Recommendations for flaterial Selection Table 17 gives tentative reconmendations for the selection of materials for CRBRP components. These recommendations are based primarily on metallurgical considerations. If design calculations indicate that the highest-strength materials are required, then H grades of' steel may be selected in order to ensure that purchased material does not contain less than 0.04 weioht percent carbon. Nevertheless, whether the normal _or H grade of steel is ordered it is recommended that the acceptable rance of carbon concentration be specified in the ordering data. The use of low-carbon rades of stainless steel are restricted to temperatures y 800 F(2 and, therefore, cannot be used in hot-leg copenents in tha coolant boundary. Their lower strength also limits their usefulness in cold-leo components. In certain situations, where there is a cost-incentive, 2-1/4 Cr-1 Mo may be used in lower temperature coolant-boundary areas in place of stainless steel. 35
A
- 7. REFERENCES
- 1. P. Soo and W. H. Horton, "The Effect of Carbon and Nitrogen ~ on the Short-Term Tensile Behavior of Solution-Treated Types 304 and 316 Stainless Steels", Topical Report, WARD-NA-3045-2, July,1973.
- 2. ASME Boiler and Pressure Vessel Code, Section II, Part A (Ferrous Materials), 1971 Edition.
- 3. Interpretations of ASME Boiler and Pressure Vessel Code, Code Case 1331-8, November 3, 1972.
- 4. B. Weiss and R. Stickler, " Phase Instabilities During High Temperature ,
Exposure of Austenitic Stainless Steel". Metallurgical Transactions, _3_ , 851 (1972).
- 5. A Thorley, B. Longson and J. Prescott, "Effect of Exposure to Sodium on the Mechanical Properties of Some Ferritic, Austenitic and High Nickel Alloys", TRG Report 1909 (C), 1969.
- 6. L. H. Kirschner, R. M. Hiltz and S. J. Rodgers, "Effect of High Temperature Sodium in the Mechanical Properties of Candidate Alloys for the LMFBR Program", MSAR 70-76, May, 1970.
- 7. G. V. Smith, An Evaluation of the Yie_ld, Tensile, Creep and Rupture Strengths of Wrought 304, 316, 321 and_347 Stainless Steels at Elevated Temperatures, ASTM Data Series Publication DS 5S2, American Society for Testing and Materials, February, 1969.
- 8. Memorandum, P. Soo to H. B. Ketchum, " Materials Data Imput into Pro-posed Sodium Component Design Guide", ARD-M&PE-PS-72-19, November 2, 1972, also reported in " Interstitial Transfer Program Impact Assess-ment Report, Part II", Topical Report by S. A. Shiels, S. L. Schrock and L. L. France, WARD-NA-3045-3, September,1973.
36
O
- 9. Memorandum, _ R. Hundal and F.1. Flagella to W. E. Ray, " Interstitial Effects on Strength of- 316 55", ARD-M&PE-PE-RH-73-01, January 19, 1973, also reported in " Interstitial Transfer Program Impact Assessment Peport, Part II", Topical Report by S. A. Shiels, S. L. Schrock and L. L. France, WARD-NA-3045-3, September, 1973.
l
- 10. C. F. Cheng, C. Y. Cheng, D. R. Diercks and R. W. Weeks, " Low-Cycle f; Fatigue Behavior of Types 304 and 316 Stainless Steel at LMFBR Operating ;
l Temperatures", from Fatigue at Elevated Temperatures, ASTM Special Technical Publication STP 520, American Society for Testing and Materials, August,1973.
- 11. J. H. Driver, "The Effect of Boundary Precipitates on the High-Tem-perature Fatigue Strength of Austenitic Stainless Steels", Metal Science Journal, 1, 47 (1971).
- 12. J. T. Barnsby and F. M. Peace, "The Effect of Carbides on the Fatigue Resistance of an Austenitic Steel", Acta Metallurgica, H , 1351 (1971).
- 13. P. Soo (compiler), " Analysis of Structural Materials for LMFBR Coolant-Boundary Components - Materials Property Evaluations", WARD-3045T3-5, November, 1972.
- 14. E. C. Bain and R. H. Aborn, "The Chromium Depletion Theory", Trans.
American Society Steel Treatment", H , 837 (1930).
- 15. ASTM-A262-70, " Recommended Practices for Detecting Susceptibility to Intergranular Attack in Stainless Steels", Annual Book of ASTM Standards, Part 3,1971, p.181.
- 16. C. M. Rosendahl, " Welding of Austenitic Stainless Steel", Svet-Saren Esab, 5, No. 1-2, 1969.
- 17. C. L. Cole and J. D. Jones, " Stress Relief of Austenitic Stainless Steels",
Stainless Steels, Iron and Steel Inst. , London,1969, p. 71. 37 _ _ . .~ _ -.- , _ _ _ _ _ _
- 18. W. L. Fleirchmann, " Heat Treatment of Welded Structures for Relief of Residual Stresses with Particular Reference to Type 347 Stainless Steel
_Weldments", Trans. ASME, 16, 6 645 (1954).
- 19. R. A. Husby, " Stress Relief of Austenitic Stainless Steels and the Associated Metallurgy", Welding Journal, 37, 3045 (1958).
- 20. W. S.' Hazelton, " Topical Report, Sensitized Stainless Steel in Westinghouse PWR Nuclear Steam Supply Systems", WCAP-7735, August,1971.
- 21. R. B. Gunia and G. R. Woodrow, " Nitrogen _ Improves the Properties of Chromium - Nickel Stainless Steels", Journal of Metals, 5_, 413 (1970).
- 22. Jones and Laughlin data sheet - Type 304-N Stainless Steel, Jones and Laughlin Steel Corporation, Stainless and Strip Division, Warren, Michigan.
- 23. L. Colombier,'" Molybdenum in Stainless Steels and Alloys", Climax Molbdenum Company Ltd., London, 1967.
- 24. M. F. Ebling and M. R. Scheil, " Time-Temperature-Sensitization (TTS)
Diagrams for Types 347, 304 L and 316 L Stainless Steels", ASTM Special Technical Publication No. 369, American Society for Testing and Materials, 1965, p. 275.
- 25. ASM Metals Handbook Vol.__l, " Properties and Selection", American Society for Metals,1964, p. 564.
- 26. R_eport on Stress Corrosion Cracking of Austenitic Chromium-Nickel Stainless Steels, ASTM STP No. 264, 1960.
D
- 27. F. L. LaQue and H. R. Copson (Edit.), C_orrosion Resistance of Metals and Alloys, Rheinhold, 1963.
l- 28. C. Edeleanu, "Transgranular Stress Corrosion in Chromium-Nickel Stainless ! Steels", J. Iron and Steel Inst., _173,140 (1953). 38
- 29. H. R. Copson, "Effect of Co :csition on Stress Corrosion Cracking of Some Alloys Containing Nickel", Physical Metallurgy of Stress Corros_ ion Fracture, T. N. Rhodin (Edit.), Interscience, 1959, p. 247.
- 30. E. E. Denhard, "Effect of Composition and Heat Treatment on the Stress
- , Corrosion Cracking of Austenitic Stainless Steels", Corrosion,16, 359t (1960).
i
- 31. D. van Rooyen, "Some Aspects of Stress-Corrosion Cracking of Austenitic Stainless Steels", Proc. First Intl. Congress on Metallic Corrosion, Butterworths, 1961, p. 309.
- 32. W. E. Loginow and J. F. Bates, " Influence of Alloying Elements on the Stress-Corrosion Behavior of Austenitic Stainless Steel", Proc.
24th. Annual NACE Conference, 1968, p. 574.
- 33. H. H. Uhlig and R. A. White, "Some Metallurgical Factors Affecting Stress Corrosion Cracking of Austenitic Stainless Steels", Trans.
American Society for Metals, 52, 830 (1960).
- 34. L. F. Scharfstein and W. F. Brindley, " Chloride Stress Corrosion Cracking of Austenitic Stainless Steel - Effect of Temperature and pH", Corrosion,14, 588t (1958).
- 35. S. P. Rideout, " Stress Corrosion Cracking of Type 3C4 Stainless Steel in High Purity Heavy Water", Proc, 2nd Intl. Congress on Metallic Corrosion, NACE,1966, p.159.
- 36. J. E. Slater and R. W. Staehle, Report No. C00-2069-16 (Q-4),
3_ Ohio State University, September, 1970. t
- 37. J. P. Hammond, " Fuels and Materials Development Program Quarterly Progress Report for Period Ending September 30, 1971", P. Patriarcha (Edit.), ORNL-TM-3550, p. 136.
39
- 38. I. L. W. Wilson and R. G. Asauen, " Caustic Stress Cerrosion Cracking of Iron-Nickel-Chromium Alleys", paper 8G-22 presented at the Intl.
Conference on Stress-Corrosion Cracking and Hydrogen Embrittlement of Iron-Base Alloys, held in Firminy, France, June,1973. Conference sponsored by NACE.
- 39. A. K. Agrawal and R. W. Staehle, " Stress Corrosion Cracking of Fe-Cr-Ni
? Alloys in Caustic Environmebts", Report C00-2018-24, Ohio State University, March, 1971.
- 40. ' G. A. Whitlow, J. C. Cwynar, R. L. Miller and S. L. Schrock, " Sodium Corrosion Behavior of Alloys for Fast Reactor Applications", Proceedings of a Symposium on Chemical Aspects of Corrosion and Mass Transfer in Liquid Sodium, Detroit, October,1971, AIME.
- 41. Steels for Elevated Temperature Service, U.S. Steel Corporation, 1965, p. 42.
- 42. R. M. Brown, R. A. Rege and C. E. Spaeder, " Evaluation of Normalized and Tempered and Quenched and Tempered 6-1/4-inch Thick Plate of 2-1/4 Cr-1 Mo Steel", in He_at-Treated Steels for Elevated Temperature Service, ASME, 1966, p. 27.
- 43. R. R. Seeley and R. M. Zeisloft, "Effect of Carbon Content in High-Temperature Properties of 2-1/4 Cr-1 Mo Steels", in Fatigue at Elevated Temperatures, ASTM Special Tech. Publication No STP 520, August, 1973,
- p. 332.
- 44. A. Thorley, IAEA Special_ists Meeting on Fission and Corrosion Produce
_ Behavior in Primary Systems of LMFBR's_, Bensberg FRG, Conf. 710959, 1971.
- 45. "LMFBR Heat Enchanger Materials Development Program" Second Quarterly Report, September-November,1972, GEAP-13919-2, December, 1972.
- 46. F. Eberle and J. H. Kitterman, " Scale Formation in Superheater Alloys Exposed to High Temperature Steam", in B_ehavior in Superheater Alloys in High Temperature, High Pressure Steam, ASME, 1969, p. 69.
40
- 47. J. P. Hammond et. al., " Corr:sion of Advanced Steam Generator Alloy Weldments in 1100 and 1200*F (595 and 650 C) Steam", in P_roceedings o_f t_he 26th Conferenc_e, National Association of Corrosion Engineers, March,1970, p. 277.
- 48. J. Hoke and F. Eberle, "0xidation of Superheater Materials by High-i Temperature Steam", ASME, Paper No. 57-A-175.
- 49. F. Eberle and J. L. McCall, " Electron Microprobe Study of Scales Formed on 2-1/4 Cr-1 Mo, 5 Cr-1/2 Mo and 9 Cr-l Mo Commercial Super-heater Tubing Aftcr 6,12 and 18 Months Exposure to 2000 psi Steam of 1100 and 1200 F, Respectively", Pr_oceedings of the American Power Conference _, 1964, Volume XXVI, ASME, 1964, p. 488.
.. 50. F. Eberle, J. W. Siefert and J. H. Kitterman, " Scaling of Ferritic Superheater Steels During 36,000 Hours' Exposure in 980/1030 F Steam of 2350 psi, With Particular Respect to Scale Exfoliation Tendency",
- ibid. p. 501.
- 51. R. A. Davies, J. A. Bray and J. M. Lyons, " Effects of Water Leakage into Sodium Systems", Alkali Metal Coolants, IAEA, Vienna, 1966, p. 263.
- 52. H. V. Chamberlain, A. J. Kanamori and P. S. Lindsey, " Evaluation of Matarials Wastage Due to Reactions of Water in Sodium", APDA-227, June, 1969.
- 53. L. G. Emmer, C. D. Clauser and J. R. Low, Jr. , " Critical Literature Review of Embrittlement in 2-1/4'Cr-1 Mo Steel", Welding Research
-Council Bulletin 183, May, 1973.
f g 54. Welding Handbook, Section 4, 6th Edition,1972, p. 63.36.
- 55. " Mechanical and Physical Properties of Austenitic Chromium-Nickel Stainless Steels at Anbient Temperatures", from C_hromium-Nickel "tainless Steel Data, Section 1 Bulletin A, International Nickel Ccmpany,1965,
- p. 38.
41
r:
- 56. "B & W. Croloy Steel Pipe, Tuting and Wt:lding Fittings", Tech. Data Card 145-A, Babcock and Wilcox Cc.,
57 ' F. Buckley, M. C. Caplan, J. J. Johnson and R. P. Kent, " Steels at
-High Temperature in Steam Turbines", The Joint International Conference on Creep, Inst. of Mech. Eng. , London,1963, p. 6-67.
L 58. L. M. Wyatt and M. G. Gemmill, " Experience with Power Generating Plants and its Bearing on Future Developments", Ibid, p. 7-1.
- 59. J. H. Harlow, " Metallurgical Experience with the Eddystone 5000 psi 1200 F Unit No. 1", Ibid, p. 7-11.
- 60. G. M. Slaughter and T. R. Housley, "The Welding of Ferritic Steels to Austenitic Stainless Steels", The Welding Journal,- 43, 454-s (1964).
- 61. J. F. Eckel, " Diffusion Across Dissimilar Metal Joints", The Welding Journal, 43, 170-s (1964).
- 62. C. L. Estes and P. W. Turner, " Dilution in Multipass Welding AISI 4130 to Type 304 Stainless Steel", The Welding Journal, 43, 541-s (1964).
- 63. R. D. Wylie, " Cooperative Investigation of a New Welding Electrode for Stainless Steel", The Welding Journal, E, 426-s (1958).
- 64. ASM Metal Handbook, Volume 6, Welding and Brazing, 1971, p. 201.
- 65. J. F. Mason, Jr. " Corrosion Resistance of Stainless Steel in Aqueous Solutions", Metals Engineering Ouarterly, 8,67(1968).
- 66. A. Feduska, S. Orbon, S. L. Schrock and S. A. Shiels, " Interstitial Transfer Program Impact Assessment Report, Part I, Application of Data to FFTF Components", Topical Report, WARD-NA-3045-3, August, 1973.
- 67. S. A. Shiels, S. L. Schrock and L. L. France, " Interstitial Transfer Program-Impact Assessment Report, Part II, Program Justification and Scope", Topical Report, WARD-NA-3045-3, September, 1973.
42
- 68. P. Soo, S. A. Shiels and S. J. Orbon, "The Effects of Interstitial Transfer on Demonstration Plant Interr.ediate Heat Exchanger Tubing", transmitted by P. Soo to D. Toler under memorandum number ARD-M&PE-PS-73-34, September 28, 1973.
T 69. " Sodium Technology Program Quarterly Progress Report for Period Ending April 30, 1973", WARD-NA-3045-1. y 70. M. C. P,0wland and D. E. Plumlee, " Sodium Mass Transfer XXII, Metallurgical Examination of Test Loops", GEAP-4838.
- 71. P. W. Koch and P. J. Kovach, " Materials Examination of a Model Sodium Heated Steam Generator", BAW-1280-37, June, 1966.
I~
- 72. J. L. Krankota and J. S. Armijo, "Decarburization Kinetics of Low Alloy Ferritic Steels in Sodium", Metallurgical Transactions, _3_, 2515 (1972).
.. 73. Private communication from F. Van Woort and W. E. Stahl, U. S. Steel Corporation, Pittsburgh', October 19, 1973.
- 74. G. N. Emmanuel, W. E. Leyda and E. J. Rozic, " Versatility of 2-1/4 Cr-1 Mo as a Pressure Vessel Material", in 2-1/4 Cr-1 Mo Steel in Pressure Vessels and Piping, ASME,1971, p. 79.
- 75. J. Sannier, 0. Konovaltschikoff, D. Leclercq and R. Darras, " Compatibility of Ferritic Steels With Sodium at High Temperatures", in Effects of Environment _ on Material Properties in Nuclear Systems, Inst. of Civil Engineers, London, 1971, p. 155. l 7 176. Private communication from C. Bagnall and S. A. Shiels, WARD.
J 43
sehank ;. . . . m . .A TABLE 1 Carbon Contents for Auste nitic Stainless Steel Prodact Fonns?_ CARBON CONTENT (ppm) AS*1E ftATERIAL 304 L ard 316 L SS 304 and 316 SS 304 H and 316 H SS SPJCIFICATION
~ ~35~740; Plate, 5heet and Strip For tusion -
Welded Unfixed Pressure Vessels : 0.03 s 0.08 , SA 249; Welded Austanitic Stainless Boiler - Superheater and Heat Exchanger and Condenser s 0.03 . ( 0.08 Tubes su SA 312; Seamless and Welded Austenitic s 0.035 0.10
- 5 0.040 for ( 0.08 0.04 -
Stainless Steel Pipe small diameter pipe SA 376; Scaraless Aastenitic Stainless 0.04 - 0.10 Steel Pipe for High Temperature 4 0.08 ' Central Station Service SA 479; Heat Resisting Steel Bars and - Shapes for Boilers ar. Other Pressure 4 0.03 1 4 0.08 Vessels
- Data from Reference 2
ha , , .
-1 d TAI:LE 2 .
5 and S "Values for Austenitic Stainless Steels
- J 5 (ksi) Sm (ksi) 316 and 304 and 316 and TEMPERATURE 304 and ,
316 H 304 H 316 L 316 H 304 L 304 H 316 L (*F) 304 L 25.0 30.0 16.6 20.0 16.6 20.0 100 25.0 30.0 21.1 25.8 16.6 20.0 16.6 20.0 200 21.3 25.0 , 23.3 16.6 20.0 16.6 20.0 300 19.1 22. 5 18.9 21.4 15.7 18.7 15.5 19.2 400 17.5 20.7 17.2 19.9 14.7 17.4 14.4 17.9 500 16.1 19.4 15.9 s 18.8 13.9 16.4 13.5 17.0 on 600 15.5 18.2 15.0 18.1 13.4 15.9 12.8 16.3 700 14.9 17.7 14.3 17.6 13.0 15.1 12.3 15.8 800 14.4 16.8 13.7 13.1 17.3 14.6** 15.7** 900 13.9 16.2 17.0 14.0** 15. 5*
- 1000 13.3 15.6 12.4 13.3** 14.8**
1100 12.7** 14.6** 1200 i
- Data are from ASME Boiler and Pressure Vessel Code, Section III and ASME Code Case 1331-8, and apply to Class I components
** For Type 304 and Type 316 gr'ades only ,
TABLE 3
. Corrosives R_eported to Induce Intergranular Corrosion in Sensitized Austenitic Stainless Steel (65) 3-3
(; Acetic acid Oxalic acid a Acetic acid + salicylic acid Phenol + naphthenic acid Ammonium nitrate Phosphoric acid Ammonium sulfate Phthalic acid
- Ammonium sulfate.+ H2504 Salt spray Beet juice Sea water Calcium nitrate Silver nitrate + acetic acid
. Chromic' acid Sodium bisulfate Chromium chloride Sodium hydroxide + sodium sulfide Copper sulfate Sodium hypcchlorite Crude oil Sulfite cooling liquor Fatty acids Sulfite solution Ferric chloride Sulfite digester acid Ferric sulfate (calcium bisulfite + sulfur dioxide)
Formic acid Sulfamic acid 1 Hydrocyanic acid Sulfur dioxide (wet) Hydrocyanic acid + sulfur dioxide Sulfuric acid Hydrofluoric acid + ferric sulfate Sulfuric acid + acetic acid ' Lactic acid Sulfuric acid + copper sulfate Lactic acid + nitric acid Sulfuric acid + ferrous sulfate ., Maleic acid Sulfuric acid + methanol Nitric acid Sulfuric acid + nitric acid 7 Nitric acid + hydrochloric acid Sulfurous acid Nitric acid + hydrofluoric acid Water + starch + sulfur dioxide Water + aluminum sulfate 46
4 TAELE 4 (a) Effect of Annealing Ter.Derature on_ Reduction of Stress in Austenitic Stainless Steel C7) Annealing Temperature Effect on Stress 400 - 750 F about 5% general relief, but reduces peak stresses by up to 40% 1000 1200 F about 35% relief 1550 1650 F about 85% relief 1750 1900 F maximum relief of about 95% TABLE 4 (b) Effect of Annealing Temperature and Time on Reduction of Stress in Type 347 Pipe (18) Annealing Temperature Residual Stress (Size 5 in. 0.D., 4 in. I.D.) As Welded 15,000 to 18,500 psi on I.D. 1200 F (4 hours) 13,700 to 15,300 1200 F (12 hours) 16,000 1200 F (36 hours) 15,600 1650 F (2 hours) 0 1850 F (1 hour) 0 (Size 9-1/4 in. 0.D., 6-1/2 in. I.D.) As Welded 26,000 to 30,000 psi on I.D. 1100 F (16 hours). 20,000 1100 F (48 hours) 20,000 D~ 1100 F (72 hours) 23,000 1200 F (4 hours) 21,500 to 24,000 l !. 47
TAELE 4-(c)
'Effect of Annealing Temperature on Relief of Stress in Type 316 Cold Drawn Tubing kl9) v .
Measured Stress, Annealing Temperature Circumferential No heat treatment,1/4' hard 52,500 psi 1000 F (24' hours, air cool) 44,500 1200 F (1/2. hour, airicool) 39,100 1375 F (1/2 hour, air cool) 26,700 1375*F (1/2 hour, furnace cool) 23,100 1450 F (1/2 hour, furnace cool) 10,400 1550 F (1/2 hour,' furnace cool) 3,600 1600 F (1/2 hour, furnace cool) . 3,500 48
4.' L' qa, p ,
- , t- ;8 TABLE 5 Results of Huey Tests on Several Heats of Nitrogen All I oyed Stainless Steel (21) i PLATE C% THICKNESS CORROSION RATE ~
N% AftflEALED (IN) (mils /yr) 1 SENSITIZED i (mils /yr) 0.035 0.070 9-1/2 11.5 __ 10.6 __ 8.9 __ 8.5 __
.066 7.0 --
0.140 7-3/4 10.0 __ 12.2 --
.056 12.2 0.089 7
.063 7.8 0.110 --
6-1/16 8.5 __
.057 7.2 0.I10 , --
5-3/4 10.3 35.6 10.0 31.4 10.3 33.1
.030 11.6 --
.100 3
.025 6.0
.130 13.2 3
.070 6.0
.040 7.2 3
6.0 84.0
TABLE 6 Huey Test Data for AIS: 304 and Jones & Laughlin 304-N (0.25 Nitrogen)(22)__ c Corrosion Rate in Mils /yr 48 hour periods Grain Material Size 1 2 3 4 5 Cold Rolled Mill Annealed Type 304-N 11 8 7 7 7 7 11 8 8 8 8 9 Avg. 8 8 8 8 8 Type 304 8 14 12 13 13 13 8 12 11 11 12 12 Avg. 13 12 12 13 13 Sensitized for One Hour at 1250?F Type 304-N 11 66 128 173 196 11 72 136 179 202 11 72, 139 180 201 Avg. 70 134 177 200 Type 304 8 58 281 454 492 8 61 313 479 513 Avg. 60 297 467 503 Annealed for 15 Minutes at 2000'F Sensitized for One Hour at 1250 F Type 304-N 4 56 44 43 45 4 41 42 50 54 4 43 50 69 78 Avg. 50 45 54 58 Type 304 2 58 39 34 34 2 45 45 35 41 2 44 35 40 42 Avg. 49 40 37 39 50 A
lllll l
)
l
, 4
( l e e t S o M 1 r m C u m 4 i m / x 1 a 2
- m
. t
' c n i e t c i r 7 r e p
r
. E e t m
u m m u u L F B h g 0 i 9 m 3 i i mm A f x x T o i 6 x 6 1 e a a a n w0 /m2 1 mm o // i 5 0 0 0 7 3 3 t 1 3 5 9 8 0 0 i
. s 0 0 0 1 0 0 0 o
p m o - - - - - - - C l a C n i r o P S c M S C M i m h e C 11llll
. , , . ., , .. ~ ~
9 TABLE 8 Low Cycle Fatigue Data for 2-1/4 Cr-1 Mo Steel (43) TEST TEMPERATURE, SPECIMEN CYCLES T0 b DEGREE F NUMBER FAILUREa INDICATION OF CRACKING MATERIAL 900 3484 3460 Standard 1 3 2736 (0.12Wt.%C) 2960 3090 4 5 3226 3216 Avg 3102 1000 6 3150 3136 7 4648 4568 8 4295 9 3115 3078 Avg 3802 900 D-2-4 1680 1650 Decarourized D-2-5 2160 2175 (0.02 Wt.% C) 1869 1905 D-2-6 D-2-7 2137 Avg 1964 s 1000 D-3-1 1961 D-3-4 1340 1362 D-3-5 2221 2211 D-3-6 2061 2073 Avg 1897 1 a Failure taken as 10 percent decrease in maximum load range - b By examination of hysteresis loops n
4 TABLE 9 Test Conditions for Steam Corrosion Tests on 2-1/4 Cr-1 Mo Steam Pressure Symbol Material Condition (psi) Boiler Water Chemistry Reference r , A -*- Annealed at 1650*F 2000 N/A 46
> Annealed at 1650*F, as 2000 N/A 46 machined surface s Annealed at 1650*F, surface 2000 N/A 46 machined and then pickled
+ Annealed at 1650*F, surface
- machined, pickled and then stress relieved at 1350*F
# Probably annealed 2000 N/A 48 v Annealed 900 SiO 0.12-0.20 ppm 2
2 4 5-25 Na 50 ppm Na 50 0.0-1.0 ppm 2 3 PO 1.0-5.0 ppm 4 Nacl 3-10 pp pH = 10.1-10.7 pp 47 x Annealed 2000 N/A 49 K Annealed 2350 N/A 50
#O Annealed -
2000 N/A 48
% Annealed at 1650*F as-machined 10 02 <10 ppb SiO 2-5 ppb 2
C1 <50 ppb Fe <10 ppb
- Resistivity >l Ma pH = 9.5 46 O Annealed at 1650*F, surface 10 46 machined and pickled
? O Annealed at 1650*F, surface 10 46 > machined, pickled and stress i relieved at 1350*F -
' b Annealed at 1650*F, as- 650 46 machined surface D Annealed at 1650*F, surface 650 46 machined and pickled e Annealed at 1650*F surface 650 46 .
machined, pickled and stress i relieved at 1350 F N/A = Not available 53
TABLE 10 Estimated Steam Corrosion Allowances for Ferritic 2-1/4 Cr-1 Mo Steel t Exposure Temperature 30-Year Corrosion Allowance (*F) (mils) 900 2.9 950 4.9 9.3 1000 1050 16.3 1100 30.1 1200 199.8 l 54 l
.a v v ~ ~ .
TABLE 11 Wastage Rates in Sodium / Water Reactions (51) MATERIAL TOTAL WT. LOSS (grams) Type 316 0.015 m 2-1/4 Cr-1 Mo 0.436 Incoloy 800 0.004 30 Milliliters of water injection over 30 minutes. Sodium Temperature 500 C (930 F) 3 Pounds Sodium Charge 1 s
n . ~. ,. . .~.-n v . . . , , TABLE 12 Coefficients of Linear Thermal Expansion for Type 304 Stainless Ste_el _a_nd 2-1/4 Cr-1 Mo Steel COEFFICIENTS OF THERMAL EXPANSION,10 -67 7 ALLOY RT - 200 400 600 800 1000 __ Type 304 SS(55) 8.8 9.1 9.4 9.6 9.8 10.2 2-1/4 Cr-1 Mo(41, 56) 6.4 6.6 7.0 7.2 7.5 7.8 8 1
TABLE 13 Chemical Composition of Filler Metals for 2-1/4 Cr-1 Mo Welds COMPOSITION (WT. PERCENT) ELEMENT' INCONEL 182 T 307 SS I) 16-8-2(c) Ni Balance 9.0 - 10.5 8.0 Cr 13.0 - 17.0 19.5 - 21.5 16.0 C 4 0.10- 0.07 - 0.15 Mn. 5.0 - 9.5 3.75 - 4.75 Fe 6.0 - 10.0 Balance Balance S 5 0.015 4 0.030
< 0.030 P
Si 41.0 0.25 - 0.60 3 Cu s 0.50 - Ti < l .0 - Nb 1.0 - 2.5(a) _ Co 5 0.12 - Mo 4 0.25 2.0 N 4 0.07 i 4 i (a) Plus Ta. Ta 0.25 maximum when specified (b) Westinghouse proprietary material (c) Approximate composition for main elements (Babcock and Wilcox proprietary
. _ material) 57
w , , . _ TABLE 14 (a) D_iffusion Depths for Hot-Leg Components MIN. THICKNESS- MAX. TEMP. INTERSTITIAL DIFFUSION DISTANCE (a) t (DercentaQe of ComDonent thickness) COMPONENT (INS) ( F) 304 (0) 304 H '316 316 H Primary System Hot Leg Vessel 2.0 995 -2.5 2.6 1.7 1.7 Vessel Outlet Nozzle 0.5 995 9.8 10.2 6.7 6.7 Piping 0.5 995 9.8 10.2 6.7 6.7 Pump Suction Nozzle 0.5 995 9.8 10.2 6.7 6.7 Pump Discharge Nozzle 0.5 995 9.8 10.2 6.7 6.7 Pump Casing 2.0 995 2.5 2.6 1.7 1.7 IHX Inlet Nozzle 0.5 995 9.8 10.2 6.7 6.7 IHX Shell 1.625 995 3.0 3.1 2.1 2.1
- g; Isolation Valve Inlet Nozzle 0.5 995 9.8 10.2 6.7 6.7 Isolation Valve Outlet Nozzle 0.5 995 9.8 10.2 6.7 6.7 Isolation Valve Body 1.5 995 3.3 3.4 2.2 2.2 Intermediate System Hot Leq IHX Outlet Nozzle 0.375 936 7.2 7.7 4.5 4.5 Piping 0.5 936 5.4 5.8 3.4 3.4 Steam Generator Inlet Nozzle 0.375 936 - - - -
Steam Generator Shell 0.5 936 - - - - Isolation Valve Inlet Nozzle 0.375 936 7.2 7.7 4.5 4.5 Isolation Valve Outlet Nozzle 0.375 936 7.2 7.7 4.5 4.5 Isolation Valve Body 0.75 936 3.6 3.9 2.3 2.3 Auxiliary System , Hot-Leg Piping 0.322 995 15.2 15.9 10.3 10.3 Sample Loop Piping 0.133 995 36.8 38.4 24.8 24.8 Valves 0.216 995 22.7 23.6 15.3 15_._3 (a) The interstitial diffusion distance is arbitrarily taken to be the depth at which the (C + N) con-centration has changed from C to within 90 percent of the starting (C + N) concentration. (b) Type 304 and Type 316 are ass'umed to have 0.06 weight percent carbon and 0.06 weight percent nitrogen. Type 304 H and Type 316 H are assumed to have 0.09 weight percent carbon and 0.09 weight percent nitrogen. i k -- e
. o.-w ___
TABLE 14 (b) Diffusion Depths for Cold-Leg Components INTERSTITIAL DIFFUSION DISTANCE (a) (DercentaQe of component thickness) MIN. THICKNESS MIN. TEMP. COMP 0NENT (INS) ( F) 304(DI 304 L
~ ~
Primary SyAtem Cold Leg Vessel 2.0 730 0.5 0.3 Vessel Inlet Nozzle 0.375 730 2.5 1.6 0.375 730 2.5 1.6 IHX Outlet Nozzle 3.7 IHX Shell 1.625 730 5.9 Isolation Valve Inlet Nozzle 0.375 730 2.5 1.6 Isolation Valve Outlet Nozzle 0.375 730 2.5 1.6 Isolation Valve Body 1.0 730 1.0 0.6 Piping 0.5 730 1.9 1.2 m I_ntermediate System Cold Leg IHX Inlet Nozzle 0.375 651 1.3 1.2 IHX Shell 0.375 651 1.3 1.2 Piping 0.375 651 1.3 1.2 0.375 651 1.3 1.2 Isolation Valve Inlet Nozzle 1.2 Isolation Valve Outlet Nozzle 0.375 651 1.3 Isolation Valve Body 0.75 651 0.7 0.6 Pump Suction Nozzle 0.375 651 1.3 1.2 Pump Discharge Nozzle 0.375 651 1.3 1.2 Pump Casing 2.0 651 0.3 0.2 Steam Generator Outlet Nozzle 0.375 651 0.5 - - Steam Generator Shell 6 51 Steam Generator Tubing 0.109 651 - (a) The interstitial diffusion distance is arbitrar,ily taken to be the depth at which the (C + N) con-centration has changed from C (b) Type 304 was assumed to 06 have weight 6.to within percent carbon90and percent of the 0.06 weight starting percent (CType nitrogen. + N) concentra 304 L was assumed to have 0.02 weight percent carbon and 0.02 weight percent nitrogen.
\
l l l l TABLE 15 a Costs for Stainless Steel Plate (73) i- ) MATERIAL COST /?no LBS. ($) 304 SS 52.00 304 L SS 58.00 316 SS 79.00 316 L SS 85.00 ? b 60
m -
-r_ .e- < , , c- .u m TABLE 16 Summary of Austenitic Stainless Steel Performance Under Li1FBR Conditions PROPERTY ALLOY RATING COMMENTS Mechanical Properties 1. 316 H SS For maximum strength Type 316 grades generally are superior.
- 2. 316 SS This is especially true for creep-rupture strength. Under
- 3. 304 H SS short-term tensile testing Type 316 has superior ductility.
- 4. ~304 SS Fabrication and 1. 304 SS There may be difficulty in avoiding sensitization during Weldability 2. 316 SS stress relief and welding of high carbon grades. Type 304
- 3. 304 H SS has better weldability than Type 316.
- 4. 316 H SS Sensitization 1. 304 SS Type 304 resists sensitization more effectively than Type 316.
- 2. 316 SS High carbon grades are likely to be sensitized at some stage
- 3. 304 H SS of welding, fabrication or stress relief.
g3 4. 316 H SS Stress-Corrosion 1. 316 SS The higher strength of Type 316 appears to be an advantage Cracking 2. 316 H SS in chloride-induced cracking. Carbon is detrimental at
- 3. 304 SS high (> 0.1 wt. percent) levels. Under caustic contamina-
- 4. 301 H SS tion conditions the relative behavior of Type 304 and 316 is unknown.
Sodium Corrosion 1. 304 SS, 316 SS, There shculd be little difference in the corrosion rates of 304 H SS, 316 H SS these alloys. Interstitial 1. 316 SS Interstitial loss from hot-leg regions, and interstitial Transfer 2. 316 H SS absorption in cold-leg regions, is minimized in Type 316
- 3. 304 SS grades of steel.
- 4. 304 H SS
a . , , . - - - . . - . - O TABLE 17(a) Tentative Material Selections for CRBRP Hot-Leg Components RECOMMENDED COMPONENT MATERIAL REASON FOR SELECTION Primary System Vessel 304 SS Adequate strength, lower cost, ease of fabricability and weldability and greater resistance, compared to 316 SS, Pump 304 SS to sensitization during fabrication and heat treatnent. Isolation Valve 304 SS Vessel Outlet Nozzle 316 SS Superior strength compared to 304 SS and better resis-IHX Inlet Nozzle 316 SS tance to interstitial transfer. Piping 316 H SS Has high mechanical strength. Specification of H grade will guarantee carbon level d. 0.04 w/o. a; Intermediate System Isolation Valve 304 SS As above for primary system valve. IHX Shell 304 SS As above for primary vessel. IHX Outlet Nozzle 316 SS As above for primary vessel outlet nozzle. Type 304 SS may be used if maximum strength not mandatory. 1 Piping 316 SS The lower operating temperature, compared to the primary hot-1 leg piping, should allow regular, rather than high-carbon, material to be used. This will reduce sensitization effects, but maintain resistance to interstitial transfer. Auxiliary System ! Coolant Overflow 316 H SS The thinner walls require maximum resistance to interstitial i and Return Piping losses. H grade quarantees carbon level > 0.04 w/o. Primary Sample Loop 316 H SS As above for coolant overflow piping. Piping ! Isolation Valve 304 SS As above for primary system valve. I 1 i i
,.g, y .. . .g -y. _
h- -4 TABLE 17(b) Tentative Material Selections for CRBRP Cold-Leg Components RECOMMENDED COMPONENT MATERIAL REASON FOR SELECTION Primary System
' Vessel 304 SS See Table 17(a)
IHX Shell 304 SS Isolation Valve 304 SS Adequate strength, _ lower cost, ease of fabricability and weldability, and greater resistance, compared to 316 SS, to Vessel Inlet Nozzle 304 SS sensitization dud ng fabrication and heat treatment. IHX Outlet Nozzle 304 SS as Piping 316 SS Superior strength compared to Type 304 and 304 H (See Table
- 2) and increased resistance to carburization and nitridation effects.
1 I_ntermediate System j Isolation Valve 304 SS As above for primary system valve. j Pump 304 SS IHX Inlet Nozzle 304 SS i i Piping 316 SS* As above from primary system piping. l
- Possible use of 2-1/4 Cr-1 No if cost is less than that for Type 316 stainless steel 1
I
4 0.18 0 0.16 - OO o {=2 g -_ s'
, , 0.14 -
0 0 CD O O.12 - 0 0 CD
- 0.10 -
00 O
, a
, O m x
'i _ i 0.08 -
E-O 0.06 - 0 0 0.04 - 0 0 0.02 - O 1 ~ I ' I I 0 O 0.02 0.04 0.06 0.08 0.10 C(wt,1.) Figure 1. Nitrogen / Carbon Ratios From Type 304 Stainless Steel Literature Review 6129-1 64
l t 80 4
$';; - T{ dz 60 - g 'T l
.i e-s /
J fj N g UTS
" E w
oc N = = % =
. e@
*"* === % '%
i j W - N N, E
. 5.-
o c , 20 YS w ~~~.- --_____- - F l l I ! ! 80 TYPE 316 STAINLESS STEEL vi .
---- TYPE 304 STAINLESS ST5EL t
]o 60 -
~ ~ _ _ _ _ _ _
S w iso - A o g i i i__ l I Y ; 800, O 20 0 , 1600 600 1000 1200
/
N TEMPERATURE (OF) Figure 2. Comparison Between the Strengths and Ductilities of Solution Treated Types 304 and 316 Stainless Steel Containing 0 Weight Percent (C + N1 6129-2 65
80 q O UTS e f'., y
\ g b
i , m 60 -
; %%~ % ; ><
i sac ~~~ m * - g ' * %
$ e -
s 5>- O YS 20 - c3
~%--~ ~ _ _ _ _ _ _ _ -_ _
2 0 ' ' ' ' I 80 TYPE 316 STAINLESS STEEL 3 - - - - TYPE 304 STAINLESS STEEL 5 60 - - r_ 5m w -
' ' ~" ~ ~ ~ ~ _ _, _
a O S 20 ' ' ' ' ' O 200 160 0 600 800 1000 1200 TEMPERATURE (D) F Figure 3. Comparison Between the Strengths and Ductilities of Solution Treated Types 304 aria 316 Stainless Steel Containing 0.04 Weight Perca:t (C + N) 6129-:1 66
100 UTS O 80 - ____ _ _; g - - - ~ E '
@:.. y ,g 7~ -U- ,
.. 60 -
*g g
- j !
_~~,% % d_ %g E g i60 E a J YS
,uJ
%~__- _ _ _ _ _ _ _ _ _
l 1 l 0 80
. ~ TYPE 316 STAINLES3 STEEL m
U _-- TYPE 304 STAINLESS STEEL
= 60 t' 2 C
E y ' a 16 0 1 0 i 1 $ l l I I 20 0 200 400 600 800 1000 1200 TEMPERATURE (OF ) l Figure 4. Comparison Between the Strengths and Ductilities of Solution Treated Types 304 and 316 Stainless Steel Containing 0.08 i Weight Percent (C + Ni l 6129 I i 67
100 i - s
, O 80 j <
- UTS z
z ' * % ~ _ _ ' '- % % E @ % M N
".i s
%z N
N $ o % 3 E N, YS S ' E 20 - g l l I I O
- 60 a
w 5 i-E 40
~~~
3 TYPE 316 STAIMLESS STEEL ' ~ ~ ~ ~ ~ ~ ~ ~ w
----TYPE 304 STAINLESS STEEL 0
I I ' I I i' s 20 O 200 400 600 800 1000 1200 l L ' TEMPERATURE (OF) i t .' l Figure 5. Comparison Between the Strengths and Ductilities of Solution l Treated Types 304 and 316 Stainless Steel Containing 0.12 l Weight Percent (C + Ni l 0129-r, 68 1 1 l 1
a , mogn _. . 1 a s amend s!dl 1 - 100 E 2 o - -
? -
[ 0 il00 F o[n o gs nt O _ +
' ' ' ' ' ' ' ' ' ' I ' ' ' ' ' ' ' ' ' ' ' ' ' '
10 100 _
- Z
~
W [ - T 10500F
-g m 4 t t; - T y---
g 10 _ O' TYPE 304 STAINLESS STEEL - Q TYPE 304L STAINLESS STEEL [ V TYPE 316 STAIMLESS STEEL _ 6 TYPE 316L STAIMLESS STEEL n/ ! I I l l illi I i i l illi i l lIIlli l l I llill l I l lliff i 10 102 103 go4 105 ! RUPTURE TIME (hr) Figure 6. Stress-Rupture for Austenitic Stainless Steel Bar and Plate i
i M ot : -4 ., (" W
. . .s' I 1
]
= 100 -
7 . y - , t - 4 4 _ V 10000 F
- 1 i 10 8 I I I I til l I I I Iil l I l' I I I 18 I I I I I i11 I I I I I I 'l 100 O Z O
, 9 O y -
+
O s O O OO l l1000F OV ! 10 O Z O -
~
I - O TYPE 304 STAINLESS STEEL D TYPE 304L STAINLESS STEEL
~
U TYPE 316 STAINLESS STEEL 6 TYPE 316L STAINLESS STEEL l 1 1 I I t i11 I l l l l l ll l l I I f ,1 l t i I I l I i 11 l l l l llll 10-1 1 10 102 103 104 MINIMUMCREEPRATE(%/105 hr) Figure 7. Minimum Creep Rates for Austenitic Stainless Steel Bar an'd Plate
r 8 a T- 1 B S~ 0 l e e- e p S t s 6 s f e i . l 0 n
'+ i S
a t o 4 4 0 1 3 4 I 0 e p y
- O T O f
~ o
- O 2 1 h I
. t 0 g n
- e 0 r t
/
S 0 e 4 I 1
)
7 r u 0 t
. p t u
* (
w R _ r N u 8 o 0 + H s O
- I 3
0 C 0 1 e
- h e e* o -
6 t
- 8 v n F '
_ C-r O I 0 0 o t 0 n 0 F o . F 0 e 0 0 0 9 0 t
*I F 0 1
0 5 n 0 1 o C F 0 e 0 3 1 ! 4 0 0 C l i i a t 0 t 0 0 s _ 1 r 2 e I 0 t n 0 I Q 8 0 f o 0 2 1 t c
- - _ - e 0 f f
0 0 0 0 0 O E - 6 5 4 3 l 8 mWoN$ g=?$= mm e . C mo r u g i F 3'b 'x
~-
lll
, W t L :fr-M o E l
( 4 L y,j t o
& ~
E, c 50 - C o x l
)
' z % -
& ,.s I
i= W a w g _ 9000F ^ G
=
5 a. m w< g = ' N m l 8 x m - m -
- o e
~
h3 Il000 F - O '~ ' O N i 10 - r .j 12000F O ,~ 4 o - ! l3000 F _- - *
')-1500 0F
- 0 ' I I ' ' ' '
- O 0.02 0.04 0.06 0.08 0.10 0.I2 0.I4 0.I6 0.I8 C+N(wt.%)
Figure 9. Effect of Interstitial Content on the 10 Hour5 Rupture Strength of Type 304 Stainless Steel Bar s
Ill 8 r 1 a
. B 0 l e
e t S 6 s s 1
. e
' l 0 n i
1 - a t W S 4 6
,w e-1 N9 -
~ -
1 0
. 3 e
i 2 ' - p y . u T f 2 o 1
' . h t
0 g n a e # e r t S 0 e pO 1 ) r
' . % u 0 t t
p . g w u F No ( N R r J 0 0 0 F 0 8 0 + H u o
. ,7 0 0 '
i 0 1 5 0 C 3 g 0 g 1 0 1 e O S h t 6 0 n o , O ' 0 o
- t F
0 n 0 e 0 t 5 n e 0 1 4 o 0 C F F l O 0 0 a O 0 i D 0 t I 3 I 1 i t s 2 r 0 e I t F 0 n 0 I 0 0 f 2 4 l o t c
- - - - - e f
k O f 0 E 0 0 0 0 0 0 6 5 6 1 3 2 1 0 1 O3 z$z N e y m ye o- e a r u g i F c ht y
\
I 1 r s a i B o 1. t S s s s i e l . l ') t -. o i n _ I a t W . S 6 AJ , v i 1 3 a.
.., ._ l 0
e p L y
#io T
f 2 o l 1 h
. t 0 g n
OT e r r t S
'6 o i )
e r o% u l t
. p t u w R
( r F A F 8 N u o F 0 0 ; 0 + H 0 0 0 0 5 0 0 C 5 0 2 9 10 l 0 1 e h t s 0 p o. n o OF r F 0 t n 0 0 0 0 0 e 0 0 t 0l 3 5 n l i 1 o 4 C i 0 l 0 a i t i t s r 2 e I 0 t n 0 I f o
' t ; ~
c
- _ - - - e o f f
0 0 0 0 0 E 5 % 3 2 1 1 1 CE E ~ wEH ,8*
- o ~ e r
u g i F e l 1 l ! j1llllll
I u , i w ww.Mi v . g 100 _ t !; ; . T _ }- G _
- Argon Exposed 5000 Hours At 1325'F 10 _
~
O _ d m O Sodium Exposed At 5000 Hours At 1325'F g - y m - - I _ Z REF. 9 0 ' ' ' ' ' ' ' ' I I I I II Il I J l l I Ill i I I IIIll i I llIIIl i 10 102 103 104 105 CREEPRATE(%/105 hr) Figure 12. Effect of Sodium Exposure on the Creep Rates of Type 316 Stainless Steel at 1325 P
x Y .- < 9 M ue$ L 3 1800
!s T .
4 $ REF. 16 0.062% C - t 1600 - 0.058% C [ 0.056% C l m
&IMO 4
a p 0.052% C 5 I E I200 i H 3 i { 0.042% C 0.010% C l 1000 - i j gg i I i ItIlt I l t i I I II I I l l lIII I l l 1IiiI
- 0.1 1 10 100 1000 '
j SEMSITIZATIONTIME(hr) l Figure 13. Onset of Sensitization in 18-8 Stainless Steel as a Function of Carbon as Determined by the Strauss Test
i \ :
' 1472 TYPE 304 + Ho T -~~_ TYPE 304
,, 1292
- [/',, s N
o N' N w , 0.010 TO OVER I.0 ipy N
$ 1202 - \
E \
^
E \ REF. 23 5 H lll2 -
\ N N
N 1022
- \ N LESS THAN 0.010 ipy g "
g N N I I ' I 932 0.1 1 10 100 1000 SENSITIZATION TIME (hr) t Figure 14. Effect of Molybdenum on the Time-Temperature-Sensitization Diagram of Type 304 (0.054% C) Stainless Steel 01"U-11 77
\
e i l w J
; i )
t 4 1500 11600
- REF. 13
' ,1300 - I200 - [{ N I100 - N . g LESS THAN 0.010 ipy OVER l.0 ipy Ng E l000 (sN*s' 5 s 0.010-0.035 ipy
~
0.035-0.100 ipy 0.100-0.350 ipy 800 - 0.350-1.0 ipy $g AE: ' 700 I I I I I I ' ' I l 0 1 3.5 10 35 100 350 1000 10,000 100,000 SENSITIZATIONTIME(.hr) 1 - Figure 15. Time-Temperature-Sensitization Diagram of Type 304 L Stainless Steel (Huey Test) 6129-15 78
i 7
* :go j -
J 1600 1500 --. . REF.13 k I o
- \g OVER I.0 ipy w NN\g E 1200 -
N$s : E LESS THAN gi100 0.0i0 ipy
. _ 0.010-0.035 ipy 0.035-0.100 ioy 0.100-0.350 ipy w'%
900 0.350-1.0 ipy l I 1 l I I I I I l ! O I 3. 5 10 35 100 350 1000 10,000 100,000 SENSITIZATION TIME (hr) 1 l l l_ Figure 16. Time-Temperature-Sensitization Diagram of Type 316 L Stainless Steel (Huey Testl l 61'11-10 2 i 79 l
- =
s, N, 33! g s ,t', );-
! ) ) a :
3 H
$ N O IM oo N NF NF \ REF. 28
\
N . N
- N u N 6 -
O\ s 100 O g 9 On 5 N 2 - g o \ O o w \ g N g 10 O 18/8 STEEL \
\
618/8 Ti-STABILIZED STEEL g D l i l 1 i 0 10 - 20 30 W) 50 1 MgCl 2 CONCENTRATION (%)
)
Figure 17. Effect of MgC12 Concentration on Time to Failure 01:20-17 80
i O 1000 i O O oo M O O O O E CRACKING O g O o 0 0 O O O@ O REF. 29 e t s
)
' +48 O I 4
B O O o 100 - O 8 0 0 h 8% T O ! g y [ No CRACKING p a E 0 O b o o "! 3 em o o oo[ E
= .
10 - 8
.1 g
A
% O w ,
A DID NOT CRACK IN 30 DAYS D ' k / I I I O 20 ig) 60 80 1 NICKEL (%) Figure 18. Effect of Increasing the Nickel Content on the Susceptibility of Fe-Cr-Ni Wires in Boiling 42% MgC1,~ 61"9-18 81
i N ! 3, ,
- i >
RELATIVF STRESS CORROSION RESISTANCE OF COMMERCIAL
- TYPE STAINLESS STEELS 70 304 309 TYPE
)
60 316 310 REF. 30
- 347 314 m 347-L d 50
, TYPE a 304
% 16 0 304-L e
O - 3 30 i 20 ( mw ,, ,,, 'M un
]
10 - I t i i i l i 0 0.1 0.5 1 5 10 50 100 500 1000 FRACTURE TIME (hr) i 3
]
4 Figure 19. Composite Curves Illustrating the Relative Stress Corrosion Cracking Resistance for Commercial Stainless Steels in Boiling 42 Percent Magnesium Chloride 61*.29-10 82
320 e 20; Mi. 16% Cr, 1.5% Mo - 40,000 LB/IM.2
--- 0 205 N i, 16% Cr, 1.5% Mo - 50,000 LB/IN.2 0 20:I Ni, 16% Cr, 40,000 LB/IN.2
- . - O 145 Mi, 16% Cr, 60,000 LB/lN.2 280 -
n h REF. 31 O i - J d
, 240 4
i i i 200 T.c h
~
1 E o ' R S a: 160 2 l W f p 120 - j O I !
/
80 - /O
. /
/
/
' /
40 - / l Y /
/
\ >
, _ __ _ _ 7 J/
0 0- 0.I 0.2 0.3 0.4 CARBONCONTENT(%) Figure 20. Effect of Carbon on Times to Failure of Stainless Steel in Boiling MgC12 0 1 29 '.x) 83
.~- .-. . _~ . - _ . _. - .__ . _ _ _ _ _ _ _ _ _ _ _ _ _
120 O 100 5 i]
~
E '
! 80
- +
[ 3 d N A to E O ANNEALED (Ref. 41) 60 0 6 NORMALIZED AND TEMPERED AT 1300 F (Ref. 42) 0 O QUENCHED AND TEMPERED AT ltS0 F (Ref. 42) i I I I I g ' h
- 80 s- &
'5 5
60 m 3 w 14 0 I
' ' I i 20 0 200 W0 600 800 1000 1200 TEMPERATURE (OF) ,
1 Figure 21. Effect of Heat Treatment on the Yield and Tensile Strengths of 2 1/4 Cr-1 Mo 61211-:.!1 84
i 1 i 1 T l } , l 30 - 3
- i E
p A $ 5 d 4 a a
' g O ANNEALED (Ref. 41) 0 i
a NORMAll2ED AND TEMPERED AT 1300 F (Ref. 42) I 0 D QUENCHED AND TEMPERED AT ll60 F (Ref. 42) I I I I I l 10 . 1 0 200 100 0 600 800 1000 1200
, \
, TEMPERATURE ('F) r i
l Figure 22. Effect of Heat Treatment on the Total Elongation of 21/4 Cr-1 Mo Steel 6129-22 85
0
.f TEMPERATURE (0) 0 0 100 200 300 400 500 600 700 60 1 I i i i i i l I 1
- 80 STANDARD 14ATERI AL REF. 43 ,
i i ('43, Po in t s)
\ -
50 70
\
\
N
\ \'
60 -a '& - 40 s E r7 50 - g 30 k %*<p-
$ 40 -
E E
~% b TENSILE y
-_t M 30 -
g f 8 f- % - 20 I STANDARD MATERIAL YlELD 20 - M 10 10 - A COMERCIAL HEAT 1893 - 0.020% C CAST TO LOW CARBON O COMMERCIAL HEAT 47963 - 0.021% C DECARBURIZED -
! 0 200 400 600 800 1000 1200 1400 TESTTEMPERATURE(OF)
Figure 23. Yield and Tensile Strengths of Low-Carbon 2 1/4 Cr-1 Mo Steels 0129-2:1 86
6 30 y r ANNEALED l4 I
--- NORM. & TEMPERED
\ *
+
REF. 74 d \
\
\ 1000 HR$
\
20 -
\\
i= \
. \
I g \
- \ 10,000 HRS
= \
d 0 \ p \ c \
\
\
10 1000 HRS
\
\ N N
100,000 HPS N 10,000 HRS N % 100,000 HRS
' I f 0 l000 l100 1200 l' TEMPERATURE (OF) i l
Figure 24. Summary of the Creep Rupture Properties of Annealed and Normalized and Tempered 2.25 Cr-1 Mo Material Ol?J J I 87
t 100 - - 80 - 60 60 - 40 0 0 gg - 1100 F (593 C) 20 REF. 43 c7 3- , ~. - - 20 - E 10 at
'.6 y-
^
0.13 C
~
g 7 12 10 -
\ 0 18 C -
6 S' 8 m > :; y;w -2 1 0.06: C > m r
,g 6 g g' n -
0.027 C g g _ m 2 2 - I i i i i tili i i t 11 i i i i I II It1 0 100 1000 10,000 100,000 RUPTURE TIME (hr) Figure 25. Stress-Rupture Curves for Laboratory lieats of 21/4 Cr-1 Mo Steels with Different Carbon Contents
'3 16 6
m 10 5 STANDARD MATERIAL - 12 - 8f 8 e* . _ = ci 5 8 - 6 -I m E e
= . -
g -e t REF. 43 $ E y - g DECARBURIZED - 2 m 3 t 1 E 0
' ' I ' '
*O a 0 .04 .08 .12 .16 .20 .24 CARB0H(wt,%)
Figure 26. Effect of Carbon Content on the Rupture Strength of 21/4 Cr-1 Mo Steel at 11000 F (593 C) 61',.")-271, 26 88
O y- . l s ,
- h '
) ' 5 x .'
TEMPERATURE (OF) i400 1300 1200 1100 1000 900 800 10 l l l l l l N O SANNIER - REF. "14 N O THORLEY - REF. M k 1 - 22 ppm 0 G.E.
* % % 25 ppm 0 RUN 8A
$ Q %'
% 0.I -
g 20 ppm 0 E
/ 3 10-2
' 6 --d 'M _ _ ~ ---6
( 8 ~ _ ppm 0 REF. 45 I I I I I 10-3 0.9 1.0 1.1 1.2 1.3 1. 4 1.5 3
, I/T(OKx10)
Figure 27. Effect of Temperature and Oxygen Contamination on Corrosion Rate of 2-1/4 Cr-1 Mo Ferritic Steel 0129-27 89
~4 T y} ,. \, } ',
' t, i
- ) }l t ! ! ,
0.05 0 2 1/4 cr-i No o
- 0.04 -
O I
?--
0.03 - 6 N E o 304 SS
= 0.02 -
2 E o E 8 0.01 - o O o I I I I I I 0 O 1000 2000 3000 16000 5000 6000 7000 EXPOSURE TIME (hr) i
,s
. > .o
+4 I Figure 28. Corrosion Rates for 2 1/4 Cr-1 Mo and Type 304 Stainless Steel in Liquid Sodium at 1200 F 0129-28 90
~
/ \ i. lj..c. d W 'I _
J E Y N 6 ,
}
/,
. /
_ / 1100 F ORNL DATA
/
/ 11000F ASME DATA 4
g - 7 x
= /
5 / x x e -.- m 3
- 7
- 36,000 HR 3 3 e-g '---
g e goooo r W 1 E 2 _g____________________ 9500 F (Predicted Line) 0
/+ ' ' ' ' ' ' ' '
O 2000 4000 6000 8000 10,000 12,000 14,000 16,000 18,000 EXPOSURE TIME (hr) Figure 29. Corrosion of 2 1/4 Cr-1 Mo in Superheated Steam at 950, 1000 and 1100 0F I
p . , 1 .~.T k s
.s/
e M 8 ..j b h~ 30 ,#
/ ORML DATA
/
25 - 20 -
- - +
C S E 15 - + e " o N CED A e-h - ASME DATA 10 5 4 t l I l I I I O 2000 14000 6000 8000 10,000 12,000 116,000 16,000 18,000 EXPOSURETIME(hr) Figure 30. Corrosion of 2 1/4 Cr-1 Mo in Superheated Steam at 12000F
o TYPE 304 STAINLESS STEEL WELDED TO 2.25 Cr. I No STEEL; NICKEL ALLOY FILLER METAL (ENiCrFe-3) Ah
- \;b
- ~,
, IE.
) ,. 5 x 1., k3 E J WELD 95 CD 620 1,. WELD METAL 580 U 20 _ 19 h \\ 17 16 15 12 13 I IN. A U II
/c s x x xn x ,
L a4 s -- --- .- i + _l_ 37 32 16 ~ 32 t BEFORE WELDING AFTER WELDING y Y SECTION A-A CONDITIONS FOR SHIELDED METAL-ARC WELDING J O I N T TY P E . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . B U TT W ELD TY P E . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . S I N GL E- U GRO O V E WELD I N G PO S IT 10 N. . . . . . . . . . . . . . . . . . . . . . . . . . . . . HO R I ZONTAL RO LLED NUMBER OF PASSES........................................... 22 0 PREHE AT AND I NTERPASS TEMPER ATUR E. . . . . . . . . . . . . . . . . . . . 500 F( a ) P0STHEAT............................................ 1350 F(b) ELECTRODE WIRE........................... 1/8 AND 5/32-IN.-DIA ENicrFe-3(c) PO WER SUPPLY. . . . . . . . . . . . . . . . . . . . . . . . . . 300- AMP MOTO R-GEN ER ATO R , CURRENT (dcrp) AND VOLTAGE: TACK WELDING, AND PASS 1...................... 60 ANP,,21 Y PASSES 2 AND 3................................ 90 AMP, 23 V P A S S E S 4 TO 22. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 120 AMP, 24 Y (a) ONLY 0 THE FERRITIC LOW-ALLOY STEEL SIDE. NEATING WAS BY INDUCTION. (b) HEATE 0 1350 F FOR I HOUR BY INDUCTION AND COOLED TO 300 F IN STILL AIR. (c) l/8-IN.-DI AMETER WIRE FOR PASSES I TO 3, AND 5/32-IN.-DI AMETER WIRE FOR PASSES 4 TO 24 Figure 31. Assembly of a Low-Alloy Steel Pipe and a Stainless Steel Pipe that was Joined with Crack-Free Welds by use of a Nickel- Alloy Filler 612ti-:11 93
e 316 H SS 304 H SS 1600 ka .
~
c
- i. _
T , s
),
1200
g 316 SS , -
/
/
7 1000 -
/ ; $ )
l
- . / .
* + I /
o no -
/
l/
} /
I / 600 -8
/
/
, 1/
400 / - t . 1
, C, FOR TYPES 304 AND 304 H SS 200 l C, FOR TYPES 316 AND 316 H SS l
I I i I 0 O 0.02 0.04 0.06 0.08 0.10 DISTANCE FROM SODitM INTERFACE (in. )
~
i l Figure 32. Interstitial Gradients in Primary System Hot Leg Components 6129-!!2 94 L
v 1600 C, FOR 304 AND 304 L SS
] le 1 \
) ',
1200 304 ss
-{
\
\
1000 g
- \
~
i \ 800 - 1
. =c I
+ \
a ; 600 -
\
\
\
. _ w _ _ _ _ _ ._ _ _ 30 t 3 3 200 -
i O I I I I
-~
0 0.01 0.02 0.03 0 . 0 11 0.05 DISTANCE FROM SODIUM INTERFACE (in.) Figure 33. Interstitial Gradients in Primary System Cold Leg Components 0129-:1:1 95
o a 1800 f 316 H SS 304 H SS 1600 I
.3
, 1400 1
l : I o 316 SS 304 SS 1200 - f
* ,7 ,
/,s 7 j /
& l /
t j
= 1000 j
+
~
o /
; /
~
/
800 - r/
/
. /
//
600 - .'/ I 140 0 ,e t C, FOR TYPES 304 AND 304 H SS C, FOR TYPES 316 AND 316 H SS
' I I l
.y 200 . 0 0.02 0.04 0.06 0.08 0.10 s DISTANCE FROM S0DIUM INTERFACE (in.) m Figure 34. Interstitial Gradients in Intermediate System liot Leg Components G129-:11 96
y e
!. o 1800 j TO 2500 ppm 1600 - } I .
b ~
')
- i d) 1400 ')
) ;
- I 1200 - 304 ss 6
e E
'(
, 1000 -
+
o 800 - 600 - l l 400 - 304 L ss l I I I I I 200 O 0.01 0.02 0.03 0.04 0.05 DISTANCE FROM SODIUM INTERFACE (in. ) I-l t Figure 35. Interstitial Gradients in Intermediate System Cold leg Components 6129-:n i 97 1 J}}