https://ntrs.nasa.gov/search.jsp?R=19710026320 2020-03-11T22:22:14+00:00Z This report was prepared as an account of Government sponsored work. Neither the United States, nor the National Aeronautics and Space Administration (NASA), nor any person acting on behalf of NASA: A.) Makes any warranty or representation, expressed or implied, with respect to the accuracy, completeness, or usefulness of the information contained in this report, or that the use of any information, apparatus, method, or process disclosed in this report may not infringe pridately owned sights; or 5.) Assumes any liabilities with respect to the use of, or for damages resulting from the use of any information, apparatus, method or process disclosed in this report. As used above, "person acting on behalf of NASA" includes any employee or contractor of NASA, or employee of such contractor, to the extent that such em+csyee or contractor of NASA, or employee of such contractor prepares, dis~eminates, or provides access to, any information pursuant to his emplcayment or contract with NASA, or his employment with such concractos. Requests for copies of this report should be referred to: Na"ii3nal Aeronautics and Space Administration Scierrtific and Technical Information Division Attention: USS-A Washington, D.C. 28546 I Figure SNAP-8 Ref rac tory B o i l e r Development F a c i l i t y Showing Mercury Dump Tank, wi th Dump Valves , I 6- EM Pump S t a t o r and Sampling Line. (C67112008) { '-- Figure 10. SNAP-8 S i n g l e Tube Tes t Bo i l e r . C68011583 TOPICAL REPORT NO. 7 SNAP-8 REFRACTORY BOILER nEVELOPMENT PROGRAM ANALYSIS AND TESTING OF A SINGLE TUBE MERCURY BOILER prepared by E. S . H s i a a p p r o v e d by R, D. B r o o k s NUCLEAR SYSTEMS PROGRAMS SPACE D I V I S I O N GENERAL ELECTRIC COMPANY C i n c i n n a t i , O h i o 45215 S e p t e m b e r 1971 p r e p a r e d f o r NATIONAL AERONAUTICS AND SPACE ADMINISTRATION NASA L e w i s R e s e a r c h C e n t e r C o n t r a c t NAS 3-10610 E d w a r d R. F u r m a n , P r o j e c t Manager TABLE OF CONTEETS Page iii. FOREWORD LIST OF ILL,USTL4TIO?;S LIST OF TABLES x i i xiii I INTRODUCTION I1 TEST APPARATUS A. Description of the Test Loop B. Description of the Test Boiler 111 INSTRUMENTATION AND OPEUTING PROCEDURES A. Loop Instrumentation B. Boiler Instrumentation 6. Loop Checkout and Operating Procedures EWPERLMENTAL RESL'LTS AND DISCUSSIONS A. Overall and Average Test Results B. Local Test Results 1, Neat Flux Distribution, Quality and Overall Heat Transfer Coefficient 2. Local Heat Transfer Coefficients an@ Correlations 3. Local Two-Phase Pressure Drop C. Discussions V CONCLUDING R-S APPENDIX I CALCULATIONPLL PROCEDURES HPEmIX I1 TABULATION OF EDERIMENTa DATA AFPENBIX III SINGLE TUBE BOILER DATA REDUCTION RESILTS FOREWORD The work described i n t h i s r epor t is p a r t of the SNAP-8 Refractory Metal Boi ler Development Program, NASA Contract NAS 3-10610, conducted by the General E l e c t r i c Company. The work was done under t h e t echn ica l management of R.D. Brooks of the General E l e c t r i c Company and E. R. Fuman of the Lewis Research Center, NASA. Pr inc ipa l t echn ica l cont r ibut ion t o t h e s i n g l e tube mercury bo i l ing program, wi th in the General E l e c t r i c Company, was done by the following individuals : D r . J.R. Peterson, Manager of Heat Transfer Engineering, provided constant guidance and h e l p f u l d iscuss ions i n the hea t t r a n s f e r ana lys i s ; M r . R.A. F u l l e r contr ibuted the t e s t i n g of t h e b o i l e r and the design of the mercury h e l i c a l EM pump; M r . J .L. Pr ide made the da ta reduction computer program; and M r . W.H. Bennethum helped i n the loop instrumentat ions. iii r7.q * n e r e s u l t s of an e x p e r i m e n t a l i n v e s t i g a t i o n of h e a t t r a n s f e r and f l u i d f l o v d u r i n g forced convec t ion b o i l i n g of we t ted mercury i n a s i n g l e t a n t a 1 . a t u b e w i t h composite v o r t e x g e n e r a t o r i n s e r t a r e p r e s e n t e d . The e x p e r i c e n t s were conducted w i t h a SaK-heated and h igh ly- ins t rumented t e s t 0 s e c t i o n at mercury s a t u r a t i o n t e a p e r a t u r e up t o 1130 F. E x p l o r a t o r y l o c a l t;qo-phase f o r c e d convec t ion b o i l i n g r e s u l t s f o r mercury a r e pre- s e n t e d f o r n u c l e a t e b o i l i n g , t h e c r i t i c a l h e a t f l u x c o n d i t i o n , t r a n s i t i o n b o i l i n g and superhea ted vapor r e g i o n s i n a once-through b o i l i n g prscess , The e x p e r i a e n t a l d a t a and the a s s o c i a t e d c o r r e l a t i o n s a r e directly a p p l i - c a b l e f o r t h e d e s i g n and o p t i m i z a t i o n of once-through b o i l e r s fcr :pace Power ?ank ine S y c l e Systems employing mercury a s t h e working f l u i d , LIST OF ILLUSTRATIONS Figure No . Page No . 1 500 KW SNAP-8 Refrac tory Bo i l e r Tes t Loop Model S ingle Tube Bo i l e r Located a t t h e P o s i t i o n Shown f o r SNAP-8 Test Bo i l e r ) . . . . . . . . . . . . . . . . . . 6 2 Three-Stage Mercury H e l i c a l Induct ion EM Pump . . . . . 8 3 Three Stage Duct Mercury EM Pump Tes t Data . . . . . . 9 4 Schematic of F a c i l i t y Gas F i red NaK Heater . . . . . . 10 5 SNAP-8 Refrac tory Bo i l e r Development F a c i l i t y Showing Mercury Dump Tank. w i th Dump Valves. EM Pump S t a t o r . . . . . . . . . . . . . . . . . . . a n d S a m p l i n g L i n e 13 6 NaK Loop During F i l l i n g Operation . . . . . . . . . . . 14 7 Photographic View of t h e S ing le Tube Tes t Bo i l e r In- s t a l l e d i n t h e 500 KW SNAP-8 Refrac tory Bo i l e r Develop- ment F a c i l i t y Loop . . . . . . . . . . . . . . . . . . 15 . . . . . 8 General E l e c t r i c S ing le Tube S t r a i g h t Bo i l e r 1 7 9 Cross Sec t iona l View of t h e S ing le Tube Tes t Bo i l e r . . 19 10 SNAP-8 S ing le Tube Tes t Bo i l e r . . . . . . . . . . . . 20 11 H e l i c a l I n s e r t . P/D = 2.0. 114-inch Center Body . . . . 22 12 Photographic View of t h e Tes t Sec t ion I n l e t Po r t ion (Exposed Before I n s u l a t i o n ) . . . . . . . . . . . . . . 23 1 3 Photographic View of t h e Tes t Sec t ion Ou t l e t P o r t i o n . (Exposed Before I n s u l a t i o n ) . . . . . . . . . . . . . . 24 14 Schematic of 550 KW SNAP-8 Refrac tory Bo i l e r Tes t F a c i l i t y wi th Ins t rumenta t ion . . . . . . . . . . . . . 27 . . . . . 15 Mercury Liquid Venturi Flowmeter C a l i b r a t i o n 29 16 NaK EM Flowmeter C a l i b r a t i o n . . . . . . . . . . . . . 31 17 C a l i b r a t i o n s of Pressure Transducers . . . . . . . . . 35 18 Ins t rumenta t ion of S ing le Tube Tes t Bo i l e r . . . . . . 36 . . . . . . . . . . . . . 19 Test Sec t ion T o t a l Heat Loss 4 1 . . . 20 Conceptual I l l u s t r a t i o n of Once-Through Boil ing. 44 LIST OF ILLUSTRATIONS (Cont'd) F i g u r e No. Page No. 2 1 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e . . . . . . . . . . . . . . . . . . . . . . P r o f i l e . . 90 22 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e . . . . . . . . . . . . . . . . . . . . . . . P r o f i l e . 9 1 23 SNAP-8 S i n g l e Tube R o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 9 2 24 SNAP-8 S i n g l e Tube R o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 93 25 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 94 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 95 SNAP-8 S i n g l e Tube B o i l e r Temperature' and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 96 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 97 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e . . . . . . . . . . . . . . . . . . . . . . P r o f i l e . . 98 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e . . . . . . . . . . . . . . . . . . . . . . P r o f i l e . . 99 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 100 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 1 0 1 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e , . , . . . . . . . . . . . . . . . . . . . . . 102 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 103 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . . . . . 104 SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . . . . . . . . . . . . . . . . . . . . 105 Comparison o f Two-Phase P r e s s u r e Drop Measured by In- . . . . . . . . . sert Thermocouples and Tay lor Gauges 106 LIST OF ILLUSTRATIONS (Cont'd) Figure No. Page No. 38 Two-Phase Pressure Drop Resu l t s Obtained from a 0.67- inch I D Tube wi th H e l i c a l I n s e r t . . . . . . . . . . . 107 Comparison of Mercury Single-Phase Vapor F r i c t i o n a l P re s su re Drop Between Tes t Resu l t s and Predic ted V a l u e s . . . . . . . . . . . . . . . . . . . . . . . . 108 . . . . Mercury S ing le Tube Tes t P re s su re Drop Resu l t s 109 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 110 Local Values of Heat Flux and Overa l l Heat T r a n s f e r . . Coef f i c i en t f o r Mercury Once-Through Boil ing Runs. 111 Local Values of Heat Flux and Overa l l Heat T rans fe r . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 112 Local Values of Heat Flux and Overa l l Heat T rans fe r . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 113 Local Values of Heat Flux and Overa l l Heat T rans fe r Coe f f i c i en t f o r Mercury Once-Through Roi l ing Runs. . , 114 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boil ing Runs. 115 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 116 48 Local Values of Heat Flux and Overa l l Heat T rans fe r . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 117 49 Local Values of Heat Flux and Overa l l Heat Transfer Coe f f i c i en t f o r Mercury Once-Through Boil ing Runs, . . 118 50 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 119 51 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boil ing Runs. 120 52 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boil ing Runs. 121 53 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 122 54 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 123 LIST OF ILLUSTRATIONS (Cont'd) Figure No. Page No. 55 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through Boi l ing Runs. 124 56 Local Values of Heat Flux and Overa l l Heat Transfer . . Coef f i c i en t f o r Mercury Once-Through oiling Runs. 125 57 Mercury Liquid Heat Transfer Data from 0.67-inch I D Tube wi th H e l i c a l I n s e r t (P/D = 2.0) Evaluated Assuming H e l i c a l Flow, Nussel t Number Versus P e c l e t Number. . . 126 58 Mercury Nucleate Boi l ing Data from 0.67-inch I D Tube wi th H e l i c a l I n s e r t (P/D = 2.0), hNB Versus q" . . . . 127 59 Mercury Nucleate Boi l ing Heat Transfer Data from a 0.67-inch I D Tube wi th H e l i c a l I n s e r t (P/D = 2.0) , q" Versus AT. . . . . . . . . . . . . . . . . . . . . . . 128 60 Nucleate Boi l ing Heat Transfer Data, hNR Versus q", a t Various q u a l i t i e s . . . . . . . . . . . . . . . . . 129 6 1 Nucleate Boil ing Heat Transfer Resu l t s , hNB Versus x, . . . . . . . . . . . . . . . . a t Various Heat Fluxes 130 62 Mercury Nucleate Boi l ing Heat Transfer Coe f f i c i en t a s . . . . . Affected by Local Radia l Acce lera t ion F ie ld . 131 63 Mercury Void F rac t ion a s a Function of Qual i ty f o r Various Flow Models, . . . . . . . . . . . . . . . . . 132 64 Comparison of Heat Transfer Models f o r Mercury Boil ing a t Wetted Conditions . . . . . . . . . . . . . . . . . 133 65 Comparison of Heat Transfer Models f o r Mercury Boil ing a t Wetted Conditions . . . . . . . . . . . . . . . . . 134 66 Funct iona l Rela t ion Between h /(ql ') 0.85 NB and 'sat . . . . According t o t h e Merucry Nucleate Boi l ing Data 135 0 85 67 Funct iona l Re la t ion Between h / [ (q" * 0.54)-j NB ) ('sat and G According t o t h e Mercury Nucleate Boi l ing Data . 136 . . 68 Cor re l a t ion P l o t of Mercury Nucleate Boi l ing Data. 137 69 Funct ional P l o t Between h~~ and (1 + AT() According to t h e Mercury Nucleate Boi l ing D a t a . . . . . . . . . . . . . . . . . . . . . . . . . 138 LIST OF ILLUSTRATIONS (Cont'd) Figure No. 70 Mercury C r i t i c a l Heat Flux Resul t s and Cor re l a t ions Compared wi th Potassium Resu l t s from References (5) . . . . . . . . . . . . . . . . . . . . . . . and (6) . 139 71 Mercury C r i t i c a l Heat Flux Resu l t s Obtained from a 0.67-inch I D Tube wi th H e l i c a l I n s e r t (P/D = 2.0). . . 140 7 2 Trans i t i on Boi l ing Data Cor re l a t ion P l o t . . . . . . . 141 73 Cor re l a t ion P l o t of T r a n s i t i o n Boi l ing Data by Assuming Flow Model wi th Constant Droplet Size. . . . . . . . . 1 4 2 74 Cor re l a t ion P l o t of T r a n s i t i o n Roi l ing Data by Assuming Flow Model w i th Constant Densi ty of Drople ts . . . . . 143 75 Liquid Entrainment a s Affected by Local Acce lera t ion . . . . . . . . Fie ld i n t he T r a n s i t i o n Boi l ing Region 144 76 Mercury T r a n s i t i o n Boil ing Heat Transfer Data and Corre la t ion . . . . . . . . . . . . . . . . . . . . . . 145 77 Mercury Superheated Vapor Resu l t s Obtained from a 0.67- inch I D Tube wi th H e l i c a l I n s e r t (P/D = 2.0). Evaluated f o r H e l i c a l Flow . . . . . . . . . . . . . . . . . . . 146 78 Measured Heat Transfer Parameter Versus I n s e r t Twist . . . . . . . . . . . . . . . . . . . . . . . R a t i o . . 147 79 Superheated Mercury Vapor Heat Transfer Data i n t h e . . . . . . . . . . . . . . . Wire Coi l I n s e r t Region. 148 80 Local Var i a t ion of Two-Phase P res su re Gradient Along . . . . . . . . . . . . . . . . . . t he Boi l ing Length 149 8 1 Local Var i a t ion of Two-Phase P res su re Gradient Along the Boil ing Length a s Affected by t h e Boi l ing P res su re 150 82 Local Var i a t ion of Two-Phase P res su re Gradient Along . . . t he Boi l ing Length a s Affected by Mass Veloci ty. 151 83 Local Var i a t ion of Two-Phase F r i c t i o n a l P re s su re . . . Gradient i n t h e Boil ing Region a t TSat = 1 1 0 0 ~ ~ . 152 84 Local Var i a t ion of Two-Phase F r i c t i o n a l P re s su re . . . Gradient i n t he Boi l ing Region a t T = 1110~~. 153 S a t 85 Local Var i a t ion of Two-Phase F r i c t i o n a l P re s su re . . . Gradient i n t h e Boil ing Region a t TSat = 1 1 2 0 ~ ~ . 154 LIST OF ILLUSTRATIONS (Cont'd) J:igure No. 86 Comparison of Experimental and Predic ted Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r a t T = 975'~. . , 155 S a t 87 Comparison of Experimental and P red ic t ed Two-Phase F r i c t i o n Pressure Drop M u l t i p l i e r a t T = 106o0F . . 156 S a t Comparison of Experimental and P red ic t ed Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r a t T = 1080°F . , 157 S a t Comparison of Experimental and P red ic t ed Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r a t T = 1 1 0 0 ~ ~ . . 158 S a t Comparison of Experimental and P red ic t ed Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r a t T = l l l o ° F . . 159 S a t Comparison of Experimental and P red ic t ed Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r a t T = 1120°F . . 160 Sa t Typical Data f o r Local Heat Transfer Res is tance from Tes t Run Number 14 . . . . . . . . . . . . . . . . . . 161 Local Nussel t Number Versus z/L f o r Slug Flow wi th S inusoida l Wall Heat Flux D i s t r i b u t i o n (Reference 31). 162 S h e l l S ide and Composite Wall Thermal Res is tance of t h e GE S ing le Tube Boi le r . . . . . . . . . . . . . . . 163 Diagrammatic D e f i n i t i o n of Terminology Employed i n t h e Ca lcu la t iona l Procedures f o r Counter-Current Once- T h r o u g h M e r c u r y B o i l e r . . . . . . . . . . . . . . . . 194 96 F r i c t i o n Fac tor Versus Axial Flow Reynolds Number f o r . . . . Tubes Containing I n s e r t s wi th Forced A i r Flow. 195 9 7 Measured Pressure Drop Parameter Versus I n s e r t Twist R a t i o . . . . . . . . . . . . . . . . . . . . . . . . . 196 98 In t eg ra t ed Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r f o r Mercury from the Modified Mart inel l i -Nelson Model. 197 99 I n t e g r a t e d Two-Phase F r i c t i o n P res su re Drop M u l t i p l i e r . . . . . f o r Mercury from the Homogeneous Flow Model. 198 100 Rat io of Two-Phase F r i c t i o n P res su re Gradient t o Liquid-Phase Gradient a t t he Same Mass Veloc i ty f o r Water a t Various P re s su res and Q u a l i t i e s (Reference 34) 199 101 General ized Mart inel l i -Nelson Model f o r Two-Phase . . . . . . . . . F r i c t i o n a l P re s su re Drop M u l t i p l i e r . 200 LIST OF ILLUSTRATIONS (Cont'd) Figure No. 101 (Cont'd) Generalized Martinelli-Nelson Model Two-Phase Frictional Pressure Drop Multiplier for the . . . . . . . . . . . . Region of 0.5 < x < 1.0. 201 102 Two-Phase Frictional Pressure Drop Multiplier for . . . . Mercury from Modified Martinelli-Nelson Model. 202 103 Generalized Homogeneous Model Two-Phase Frictional Pressure Drop Multiplier . . . . . . . . . . . . . . . 203 104 Two-Phase Frictional Pressure Drop Multiplier for . . . . . . Mercury from Homogeneous Model (n = 0.25). 204 LIST OF TABLES Table No. 1 Loop Thermocouple Description 2 Boiler Insert Thermocouple Description 3 Eoiler Insert Thermocouple Description 4 Overall and Average Results 5 Overall Pressure Drop Results 6 Overall Pressure Drop Results in Two Phase Region 7 Subcooled Liquid Results 8 Local Nucleate Boiling Heat Transfer Results 9 Critical Heat Flux Resu l t s 10 Local Transition Boiling Heat Transfer Results 11 Local Superheated Vapor Results 12 Local Pressure Drop Data in Two-Phase Region Page No. 207 xii NOMENCLATURE Units Symbols A Area Radial Acceleration Constant Pressure Heat Capacity Diameter Diameter of Liquid Droplet o r Wire Coi l Fract ion of Entrained Liquid Darcy-Weisbach F r i c t i o n Factor Mass Velocity Grav i t a t iona l Conversion Factor Beat Transfer Coefficient Latent Heat of Vaporization Energy Conversion Factor S l i p Ratio Thermal Conductivity Heat Transfer Parameter f o r Hel ica l Flow Pressure Drop Parameter f o r Hel ica l Flow Boiler Tube Length Number of Liquid Droplets per Unit Area hD Nusselt Number, - DGCp Pecle t Number, - k PCP Prandt l Number DG Reynolds Number - lJ Pressure I n s e r t T w i s t Ratio Rate of Heat Transfer x i i i Btullb, 778-f t-lb /Btu m -- -- ps ia -- Btu/sec Symbols Units - q" Heat Flux Btu/hr-f t 2 T Temperature OF U Overall Heat Transfer Coefficient Btu/hr-f t'-@F V Velocity f t / s e c x Flowing Quali ty -- z Distance Along Boiler Tube from Mercury I n l e t f t a Void Fract ion -- A Difference -- € m Eddy Diffusion f o r Momentun 8 Elevation Angle degree Dynamic Viscosi ty Kinematic Viscosi ty lbm/f t-hr f t 2 / h r P Density l b /£t3 m In tegra ted Two-Phase F r i c t i o n a l Pressure -- Drop Mul t ip l ier 52 Entrainment Parameter Defined i n Equation ( 4 6 ) -- ? Local Two-Phase F r i c t i o n a l Pressure Drop Mul t ip l i e r -- Subscripts a Axial B Boiler o r Boiling b Bulk Fluid Temperature BN Bubble Nucleation c Refers t o C r i t i c a l Heat Flux Condition cb Refers t o I n s e r t Centerbody cw Refers t o Combined Wall e Equivalent o r e x i t o r e l eva t ion Subsc r ip t s exP Experimental Resu l t s f F r i c t i o n a l P re s su re Drop FE Refers t o Film Evaporation H Refers t o H e l i c a l Flow Condition o r Homogeneous Model Hg Mercury Flow i n Tube HV H e l i c a l Vane I n s e r t i I n s i d e o r I n l e t I B Refers t o Boi l ing Incep t ion Po in t IS Refers t o Superheat Incep t ion Po in t L Refers t o Liquid Phase LB Refers t o Liquid Boi l ing m Momentum Pres su re Drop M Martinel l i -Nelson P res su re Drop Model NaK NaK-78 Flow i n S h e l l NB Refers t o Nucleate Boi l ing Region o Outside o r Ou t l e t PB Refers t o Pool Boi l ing s S h e l l Sa t Refers t o S a t u r a t i o n Condit ion s c Refers t o Subcooled Heating Region SH Refers t o Vapor Superheat ing Region o r Degree of Superheat SP S ing le Phase 8s S t a i n l e s s S t e e l t ,T Tube o r T o t a l ta Tantalum Subscripts TB Refers to Transition Boiling Region TP Two-Phas e V Refers to Vapor Phase w Tube Wall wc Wire Coil Insert x Evaluated a t Local Quality Condition Superscripts - Averaged Value xvi I . INTRODUCTION The once-through b o i l e r holds promise of being h igh ly s u i t a b l e f o r l i q u i d metal Rankine Cycle Power Systems. I n a t y p i c a l high- performance ve r s ion of t h i s concept, l i q u i d e n t e r s t h e b o i l e r a t subcooled cond i t i ons , where it i s heated t o b o i l i n g temperature , bo i led t o d r y s a t u r a t e d vapor cond i t i ons , and then hea ted t o superheated vapor f o r d i scharge t o t h e t u r b i n e - a l l i n one cont inuous pass through t h e b o i l e r . The b o i l e r tube u s u a l l y con ta in s vor tex gene ra to r i n s e r t s t o promote b o i l i n g s t a b i l i t y and t o i n c r e a s e hea t t r a n s f e r performance. Heat ing of t h e b o i l e r is provided by a l i q u i d meta l r e a c t o r coolan t . The SNAP-8 Rankine Cycle conversion system has been udder development by NASA-LRC s i n c e 1963. Mercury is used a s t h e working f l u i d whereas t h e energy source is provided by a nuc l ea r r e a c t o r coo lan t , NaK. I n t h e development of a once-through b o i l e r f o r t h e SNAP-$ system, major problems encountered have been mercury and NaK i n c o m p a t i b i l i t y wi th t h e containment m a t e r i a l and t h e unp red i c t ab l e thermal and hydrau l i c performance ( I ) ( 2 ) . I n t h e f a l l of 1966, NASA-LRC recognized t h e neces s i t y of changing t h e b o i l e r tube m a t e r i a l and s e l e c t e d tantalum a s t h e new mercury-containment m a t e r i a l due t o i t s high r e s i s t a n c e t o m e r - cury cor ros ion . The second problem of understanding t h e n a t u r e o f b o i l i n g mercury i n a once-through b o i l e r w i th vo r t ex gene ra to r i n s e r t s was not explored u n t i l l a t e 1967. Ea r ly in 1967, a series of t e s t s were conducted a t t h e General E l e c t r i c Company on mercury-tantalum systems t o explore t h e thermal w e t t a b i l i t y of mercury i n con tac t wi th tantalum su r faces . An import ant r e s u l t (3) of t hese t e s t s was t h e high degree of w e t t a b i l i t y of tantalum exhib i ted by mercury a t e l eva t ed temperatures . I n add i t i on , t h e r e s u l t s i nd ica t ed t h a t once we t t i ng had been obtained, t h e mercury d id not become nonwetting upon coo l ing i n t h e absence of a i r . These i n i t i a l f i nd ings were l a t e r v e r i f i e d independently by con tac t angle measurements conducted by Geoscience, Ltd. (4) . Resu l t s i n Reference (4) i n d i c a t e t h a t a t temperatures g r e a t e r t han 1000°F t h e contac t angle between mercury and tantalum approaches z e r o (pe r f ec t wet t ing) . Based upon t h e s e e a r l y f i nd ings i n 1967, a dec i s ion was made t o des ign a s i n g l e tube tantalum mercury b o i l e r by us ing t h e hea t t r a n s f e r and two-phase pressure drop r e l a t i o n s h i p s f o r a we t t i ng f l u i d . These relationship^(^) (@were developed under NASA c o n t r a c t s 3-2528 and 3-9426 f o r high temperature potassium, a h igh ly we t t i ng f l u i d . Double containment u s ing s t a t i c NaK a s a b a r r i e r between t h e s h e l l s i d e primary NaK flow and t h e tube s i d e mercury was employed. The s t a t i c NaK was loca ted i n t h e annulus between an ou te r s t a i n l e s s s t e e l tube and an inne r tantalum tube , with mercury flowing i n s i d e t h e tantalum tube. A composite vortex genera tor i n s e r t , c o n s i s t i n g of h e l i c a l vane and wire c o i l s e c t i o n s , was pos i t ioned i n s i d e t h e b o i l e r t ube t o i nc rease t h e hea t t r a n s f e r capac i ty . E a r l i e r SNAP-8 b o i l e r designs were based upon nonwetting mercury h e a t t r a n s f e r relations''') and r e s u l t e d i n b o i l e r tube l eng ths of 30-feet f o r t h e SNAP-8 system requirements. The use of wet t ing hea t t r a n s f e r r e l a t i o n s i n t h e thermal design ca lcu la t ions f o r t h e s i n g l e tube b o i l e r r e su l t ed i n t h e s e l e c t i o n ~f a s h o r t e r 16-foot tube length f o r t h e same system operat ing condit ions. During t h e end of 1967 and e a r l y i n 1968, t h e s i n g l e tube mercury b o i l e r was designed, f ab r i ca ted and success fu l ly t e s t e d i n t h e 550 MW capaci ty l i q u i d metal f a c i l i t y located a t t h e Evendale Plant . Bo i l e r operat ion included 326 hours of t e s t i n g and t h r e e r e s t a r t cycles. No d e t e r i o r a t i o n i n performance was observed wi th in t h e spec i f i ed SNAP-8 power system operat ing ranges. A l a rge amount of t e s t da ta was accumulated and subsequently analyzed t o explore t h e heat t r a n s f e r and pressure loss c h a r a c t e r i s t i c s f o r t h e once-through bo i l ing of mercury i n s i d e a tantalum tube with vortex generators . Detai led descr ip t ions of t h e s i n g l e tube b o i l e r assembly, t e s t apparatus and instrumentat ion a r e presented i n subsequent sec t ions , Local and o v e r a l l boi l i n g heat t r a n s f e r r e s u l t s and t h e i r c o r r e l a t i o n s a r e presented f o r a l l 16 planned t e s t runs. Individual heat t r a n s f e r regions, designated as subcooled l iqu id , nuclea te boi l ing , t r a n s i t i o n boi l ing and superheated vapor, were i d e n t i f i e d from t h e NaK and mercury temperature p r o f i l e s es tabl i shed by a la rge number of ca l ib ra ted thermocouples i n s t a l l e d on both the s h e l l ex te rna l su r face and i n s i d e t h e b o i l e r i n s e r t . Local and overa l l pressure losses and t h e i r co r re la t ions f o r bo i l ing mercury a r e a l s o presented and compared with e s t ab l i shed predic t ions. Local two-phase pressure loss gradients , l o c a l pressure loss mul t ip l i e r s , and pressure losses f o r superheated mercury vapor i n t h e wire c o i l s e c t i o n of t h e i n s e r t were thoroughly analyzed and corre la ted . Two e x i s t i n g two-phase pressure loss c o r r e l a t i o n s , t h e M a r t i n e l l i model and homogeneous model, are compared wi th t h e experimental two-phase p re s su re l o s s r e s u l t s . Performance t e s t i n g w a s a l s o accomplished by varying t h e ope ra t ing condi t ions over t h e expected ope ra t ing band f o r t h e SNAP-8 system t o o b t a i n off-design information. Three hundred and twenty-six hours of performance t e s t i n g and t h r e e restart c y c l e s were achieved wi thout any d e t e r i o r a t i o n i n performance i n t h e s i n g l e tube mercury b o i l i n g experi-e ment. Accurate temperature p r o f i l e s of bo th s h e l l s i d e NaK flow and tube s i d e mercury flow and a l a r g e amount of t e s t d a t a were secured and subsequent ly analyzed. The experimental h e a t t r a n s f e r and p re s su re drop r e s u l t s of b o i l i n g mercury are presented i n t h e subsequent chap te r s . Complete t a b u l a t i o n s of t h e l o c a l and o v e r a l l h e a t t r a n s f e r and p re s su re l o s s e s obtained from t h e p re sen t test a r e given i n t h e appendix. The ranges of ope ra t ion f o r a l l 326 hours of t e s t and t h r e e restart cyc le s a r e l i s t e d as fol lows: Mercury Flow Rate 700 t o 1700 l b / h r NaK Flow Rate 6300 t o 7350 lb /h r Plercury Ex i t Temperature 1290 t o 1335°F NaK I n l e t Temperature 1292 t o 1 3 5 0 " ~ Mercury Ex i t P re s su re 75 t o 267 p s i a Mercury S a t u r a t i o n Temperature 975 t o 1 1 3 0 ' ~ Degree of Vapor Superheat 205 t o 390°F Thermal Power 36 t o 75 KW 11. TEST APPARATUS A . Descript ion of t h e Test Loop The NaK-mercury test f a c i l i t y employed i n t h e experiment i s shown schematical ly i n Figure (1) . The e n t i r e f a c i l i t y was constructed of Type 316 s t a i n l e s s s t e e l except t h e tantalum b o i l e r tube, t h e L-605 NaK dump tank and t h e 6 6 0 5 mercury vapor l i n e . A s shown i n t h e f igure , t h e NaK primary loop of t h e f a c i l i t y , which simulates t h e r e a c t o r coolant loop of an ac tua l SNAP-8 power conversion system, accepted heat from a 550 KW gas f i r e d furnace and re jec ted heat t o bo i l ing mercury i n t h e secondary loop. A r e v e r s i b l e h e l i c a l induction EM pump was used i n t h e s i n g l e phase primary loop and t h e NaK flow r a t e was determined with an electromagnetic flowmeter. The NaK l e v e l i n t h e dump tank was indica ted by a l e v e l probe and was cont ro l led by i n e r t gas p ressu r i za t ion . Mercury flowing i n t h e secondary loop was boiled and superheated i n s i d e t h e tantalum t e s t b o i l e r and was condensed i n an air-cooled, f inned-tube condenser. An adjus t able vapor t h r o t t l e valve posi t ioned downstream of t h e b o i l e r was employed t o maintain both b o i l e r e x i t and condenser i n l e t pressures a t the proper system values. The valve normally operated i n a choked mode. A valved mercury dump tank connected t o t h e loop downstream of t h e condenser. The l i q u i d mercury flowed through a h e l i c a l induction EM pump, a venturi flowmeter and f i n a l l y a l i q u i d t h r o t t l e valve upon its r e t u r n t o t h e test s e c t i o n from the dump tank. The mercury vapor 5 Figure 1, 500 KW SNAP-8 Refractory Boiler Test Loop Model (Single Tube Boiler Located at the Position Shown for SNAP-8 Test Boiler. (CDC11711) line was a 3 1/2-inch 0 , D . Haynes-25 seamless pipe. Since the mercury was superheated vapor and the Haynes-25 was chosen for its avai labi l i ty and high strength. The vapor throttle valve was also made of Hayes- 25 alloy. The three-s tage mercury h e l i c a l induct ion EM pump a s shown schematical ly i n Figure (2)was designed, f ab r i ca ted and t e s t e d . The pump consisted e s s e n t i a l l y of a polyphase s t a t o r , a pump duct made up of two concentr ic tubcs with s p i r a l passages i n t h e annulus and s u i t a b l e enclosing and support ing framework. A s shown i n Figure(31, t e s t i n g of t h e mercury EM pump using a three-stage pump and a 15Snch- long s t a t o r , indica ted , a t the design point , using approximately 560 v o l t s would provide 560 p s i developed head a t t h e SNAP-8 system design flow r a t e of 12,500 lb/hr . No d i f f i c u l t i e s whatsoever were encountered i n t h e operat ion of t h e h e l i c a l induction pumps f o r pumping l iqu id mercury. The gas-f ired furnace employed as t h e prime heat source i s shown schematical ly i n Figure(4) . The furnace was designed f o r a nominal heat load of 550 KW a t a NaK e x i t temperature of 1350°F. The hea te r s h e l l was a i r - t i g h t and t h e f l u e was valved s o t h a t NaK fumes could be contained i n t h e event of a leak. An a l t e r n a t e vent l i n e t o a scrubber system insured t h a t no c a u s t i c fumes would be re leased t o t h e atmosphere. Type 316 s t a i n l e s s s t e e l wxs se lec ted a s t h e mater ia l of cons t ruct ion f o r t h e air-cooled, finned-tube mercury dondenser. This selection was based on t h c welding and post-weld annealing l-l RJ v .d & +J C) (U l-l W Flow Rate, lb /hr F igure 3. Three Stage Duct Mercury EM Pump T e s t Data. Scrubber Heater S tack Connection - NaK O u t l e t Casing Tube Bundle NaK I n l e t Supports Burner Assembly F igure 4. Schematic of F a c i l i t y Gas F i r ed NaK Heater . requirements involved with the use of t h e more corrosion r e s i s t a n t low a l l o y s t e e l s and the l ack of s i g n i f i c a n t corrosion of t h e a u s t e n i t i c s t a i n l e s s steel a t t h e condenser temperature of 45Q°F t o 600 '~ . The condenser cooling a i r was drawn from outs ide the building t o t h e con- denser. Two blowers were used f o r exhausting a i r from t h e condenser t o improve the r e l i a b i l i t y of cooling a i r t o the condenser. In addi- t i o n , a screen was i n s t a l l e d over the a i r in take duct t o minimize any in take of fore ign objects . The mercury dump tank i n s t a l l a t i o n wi th some of t h e mercury piping and valves is shown i n Figure (5). The mercury cooler shown was used t o t r i m the l i q u i d temperature re turning t o t h e pump from the condenser. The mercury EM pump s t a t o r is shown i n s t a l l e d a t t h e l e f t s i d e of t h e p i c t u r e and, thus operated with a s l i g h t negative pressure i n the test cell. The a i r discharge was a l s o connected t o the scrubber piping and, i n the event of a leak, t h e a i r could be discharged t o t h e s a f e t y system. On t h e r i g h t s i d e i s shown t h e mercury sampling l i n e t o t h e sample s t a t i o n which was ex te rna l to the enclosure t o permit sampling during operat ion. Similar connections f o r sampling were i n s t a l l e d t o sample primary loop NaK and s t a t i c NaK from the b o i l e r during the t e s t . For the purpose of s a f e t y , a mercury vapor de tec to r was i n s t a l l e d i n the condenser a i r e x i t ducting. A s i m i l a r de tec to r was i n s t a l l e d t o monitor the test c e l l f o r mercury vapor. Figure (6) shows a p a r t i a l view of t h e NaK loop dur ing the NaR f i l l i n g ope ra t ion f o r t h e primary loop. I n F igure (7), w i t h i n t h e mercury enclosure , t h e mercury loop had j u s t been completed w i t h t h e s i n g l e tube t e s t b o i l e r i n s t a l l e d . The i n t e r i o r of t h i s enc losure w a s coated wi th two l a y e r s of wh i t e epoxy p a i n t t o f a c i l i t a t e clearn- up i n t h e event of a mercury s p i l l . E l r a c t r i c heat c r s were provi ded t o prehea t t h e mercury loop and t h e NaK dump t ank cluri ng t h c s t a r t u p period. The 550 KW l i q u i d metal test f a c i l i t y and i t s components a r e d iscussed i n g r e a t e r d e t a i l i n Reference (7) i n which t h e des ign and c a l c u l a t i o n s a r e a l s o presented. B. Descr ip t ion of t h e Test B o i l e r The s i n g l e t ube mercury b o i l e r is shown schemat ica l ly i n F igure (8). The b o i l e r employed a s i n g l e tube counterf low conf igu ra t ion with a s t a t i c NaK annular l a y e r a c t i n g a s a b a r r i e r between s h e l l s i d e Nag flow and tube s i d e mercury flow. The b o i l e r s h e l l and t h e s t a t i c NaK containment t ube were f a b r i c a t e d from 316 SS tubing . A c r o s s s e c t i o n a l view of t h e t e s t b o i l e r assembly i s shown i n ~ i ~ u r e ( 9 ) . The tan ta lum b o i l e r tube had an i n s i d e diameter of 0.67-inch and an o u t s i d e diameter of 0.75-inch. The s t a t i c NaK containment t ube had an i n s i d e diameter of 0.805-inch and a wa l l t h i ckness of 0.035-inch, w h i l e t h e b o i l e r s h e l l had an i n s i d e diameter of 1.38-inches and a wall t h i ckness of 0.062-inch. The o v e r a l l hea t t r a n s f e r length was 16 f e e t . A disassembled view of t h e t e s t b o i l e r i s shown i n F igu re ( l0 ) . 12 il i z d 2 .!-I 2 r S t a i n l e s s S t e e l S h e l l S t a i n l e s s S t e e l Annulus Tantalum Tube NaK Primary Flow Figure 9, Cross Sectional View of the Single Tube Test Boiler. Spacer p ins welded one f o o t a p a r t on t h e s t a t i c NaK containment tube allowed i t t o be accu ra t e ly centered i n t h e s h e l l dur ing assembly, Three b i m e t a l l i c j o i n t s between tantalum and s t a i n l e s s s t e e l were employed. One a t each end of t h e b o i l e r tube and one a t t h e i n l e t of t h e h e l i c a l vane i n s e r t . The composite i n s e r t employed i n t h e test * cons is ted of a P/D = 2.0 h e l i c a l vane i n s e r t f o r t h e f i r s t 1 2 f e e t and a P/D = 0.97 wi re c o i l i n s e r t w i t h w i r e diameter equal t o 0,094- inch i n t h e l a s t four f e e t of t h e t e s t b o i l e r . A photographic view of t h e h e l i c a l vane i n s e r t is shown i n F igure ( l l ) , It contained a l /4- inch OD, 0.05-inch t h i c k w a l l hollow centerbody and 0.035-inch t h i c k h e l i c a l vane. The hollow centerbody permi t ted l o c a l mercury temperature measurement by t h e i n s t a l l a t i o n of thermocouples i n s i d e it. The h e l i c a l vane i n s e r t was terminated be fo re t h e b o i l e r e x i t i n order t o avoid carry-over of any l i q u i d mercury which had adhered t o t he centerbody. A s t h e mercury flow entered t h e b o i l e r tube , a h e l i c a l flow p a t t e r n was induced by t h e i n s e r t . The purpose of t h e arrange- ment was t o i n c r e a s e t h e h e a t t r a n s f e r capac i ty of t h e b o i l i n g f l u i d , p a r t i c u l a r i l y i n t h e t r a n s i t i o n reg ion , s i n c e t h e r a d i a l a c c e l e r a t i o n produced by t h e i n s e r t w i l l tend t o f o r c e t h e l i q u i d d r o p l e t s en- t r a i n e d i n t h e vapor c o r e t o t h e p a r t i a l l y wet ted tube wal l . Photographic views of t h e b o i l e r i n l e t and o u t l e t po r t ion j u a t be fo re i n s t a l l i n g i n s u l a t i o n m a t e r i a l s a r e shown i n Figures (12) and (13). The s h e l l - s i d e thermocouples a r e a l s o i l l u s t r a t e d by theee f i g u r e s , As shown i n Figure (12), r e l a t i v e axpansion bstwsan t h ~ bogllsr tube and the shell of tha erse@mbly wae accomxnodr;rted by ra b~llswa positioned near the mercury inlet and of the teat rect ios , rC Sse Nomenclature r-. 0 c-4 il b F i g u r e 13. Pho tograph ic View of t h e T e s t S e c t i o n O u t l e t P o r t i o n * (Exposed Before I n s u l a t i o n ) . ((267120768) 111. INSTRUNIENTATION AND OPERATING PROCEDURES A. Loop Pnst rument a t i o n A complete ins t rumenta t ion layout f o r t h e NaK-mercury hea t t r a n s f e r loop i s shown on t h e loop schematic drawing i n F igure (14) . Pressure gauges, flowmeters, thermocouples, l e v e l c o n t r o l s and mercury vapor d e t e c t o r s a r e shown i n t h e i r approximate loca t ions . The i n s t ru- mentat i o n i n s t a l l a t i o n c o n s i s t s of temperature and p re s su re ins t ruments which monitor ope ra t ion and provide d a t a f o r h e a t t r a n s f e r and p re s su re l o s s a n a l y s i s . A l i q u i d v e n t u r i designed t o provide 20 p s i d i f f e r e n t i a l p re s su re a t 1800 l b / h r mercury flow was i n s t a l l e d i n t he loop f o r t h e mercury flow r a t e measurement. P r i o r t o i n s t a l l a t i o n water was used t o determine the v e n t u r i d i scharge c o e f f i c i e n t over app ropr i a t e t h r o a t Reynolds number range. The mercury flow r a t e was then determined by t h e s tandard v e n t u r i flow equat ion a s a func t ion of t he d i f f e r e n t i a l p re s su re ( i n l e t p ipe t o t h r o a t s t a t i c ) . An on-loop c a l i b r a t i o n of the v e n t u r i d i f f e r e n t i a l p re s su re versus t h e output s i g n a l was done i n 10% increment of t h e f u l l s c a l e (0 t o 20 p s i ) i n both d i r e c t i o n of i nc reas ing and decreas ing , A f i n a l c a l i b r a t i o n curve f o r t he l i q u i d v e n t u r i , p l o t t e d a s mercury flow r a t e versus output s i g n a l can thus be e s t a b l i s h e d and is presented i n Figure (15). Figure 14. Schematic of 550 KW SNAP-8 Refractory Boiler Test Facility with Instrumentation. A permanent magnet flowmeter was used t o measure NaK flow. Th i s meter c o n s i s t s of a 3.513-inch OD x 2.812-inch I D Haynes-25 c y l i n d r i c a l duct loca ted w i t h i n a 728 gauss magnetic f i e l d t o provide a t h e o r e t i c a l l y c a l c u l a t e d s e n s i t i v i t y of 13.93 gpm/mv a t 500°F. A c a l i b r a t i o n curve f o r NaK flowmeter is shown i n F igure (16). A t l o c a t i o n s where it was necessary t o determine f l u i d temperature , thermocouple we l l s were provided which w e r e immersed i n t h e f l u i d s t ream such t h a t t h e depth of immersion was a t l e a s t t e n t i m e s t h e wel l diameter . The w e l l s were made from 0.25-inch OD t u b i n g and 0.062-inch OD sheathed thermocouples w i th capped, non-grounded junc t ions i n s t a l l e d t h e r e i n . With two except ions , a l l thermocouples w e r e chromel-alumel t ype wi th 0.125-inch, 0.062-inch o r 0.04-inch OD. One hundred and f i f t y thermocouples were i n s t a l l e d on t h e e n t i r e test system wi th 71 on t h e s i n g l e t ube t e s t b o i l e r . A l l thermocouples necessary f o r per- formance c a l c u l a t i o n s were connected i n t o a d i g i t a l record ing system v i a a 150 OF c o n t r o l l e d tempera ture r e f e r ence j unc t ion hea t block. Two copper cons t an t an thermocouples re fe renced t o an ice ba th were used t o check t h e r e f e r ence block temperature . A l l o t h e r thermocouples were connected t o va r ious readout ins t ruments loca ted on t h e c o n t r o l console f o r monitor ing du r ing ope ra t i on . Approximately 50 of t h e thermocouples, which were r equ i r ed f o r d a t a r educ t ion , were c a l i b r a t e d p r i o r t o i n s t a l l a t i o n a t t h e f r e e z i n g p o i n t s of z inc (787. 1°F) and aluminum ( L220°F). 'l'he c r d i brat.ed thermocouples included a l l t h e we1 1 t,hcrmocouplcs and a r e p r e s e n t a t i v e grqup of t h e b o i l e r s h e l l s u r f a c e thermocouples. C a l i b r a t i o n r e s u l t s a r e summarized a s follows: Duct Output S ignal , emf, m v Figure 16. NaK EM Flowmeter Ca l ib ra t ion . 0.04-inch OD i n s e r t thermocouples Maximum c r r o r 1. 1°F w i i h most less than 0.5"F. t 0.062-inch OD we l l thermocouples Within a - 3/8% tol t>rnnce i n a l l except two cases + which were wi th in - 0.5%. 0.125-inch OD Exposed Junct ion + Within a - 3/4% t o l e r a n c e thermocouples i n a l l cases . The d i f f e r e n t i a l p re s su re gauges (P4, P7) , as shown i n F igure (14) were used t o measure p re s su re drop through t h e two mercury flow nozz les and another (P2) which w a s used t o measure s h e l l s i d e p re s su re drop f o r NaK flow a r e s l a c k diaphragm u n i t s . Each u n i t employed a 0.005- inch t h i c k diaphragm which t r ansmi t s t h e p re s su re f o r c e s t o a d i f - f e r e n t i a l sens ing chamber (mounted o u t s i d e t h e loop enc losure) v i a a NaK f i l l e d c a p i l l a r y tube. The gauges provide a pneumatic output s i g n a l which v a r i e s from 3 - 15 p s i over t h e c a l i b r a t e d range of t h e p re s su re t ransducer (0-600-inches of water f o r mercury t ransducers and 0 - 5 p s i f o r t h e NaK t ransducer ) . The 3 - 15 p s i ou tput s i g n a l is converted i n t o a 4 - 20 ma dc c u r r e n t by a s i g n a l condi t ion ing c i r c u i t i n t h e c o n t r o l room. The dc cu r r en t i s then converted i n t o a 1 0 - 50 mv s i g n a l f o r recording a t t h e c o n t r o l console and a 2 t o 10 v o l t s s i g n a l f o r d i g i t a l recording. C a l i b r a t i o n of t h e t h r e e d i f f e r e n t i a l p re s su re t r ansduce r s was accomplished p r i o r t o i n s t a l l a t i o n by s u b j e c t i n g t h e high s i d e t o a s e r i e s of i n e r t gas p re s su re s t e p s t o cover t h e c a l i b r a t i o n range with t h e low s i d e vented t o atmosphere. Th i s was done a t both ambient temperature and a t 500 O F (nominal). A l l c a l i b r a t i o n r e s u l t s showed + t h a t t h e t ransducers were l i n e a r and r epea t ab le w i th in - 1% throughout t h e c a l i b r a t i o n range. Typica l c a l i b r a t i o n curves a r e g iven i n F igure (17 ) . F i n a l z e r o adjustments were made wi th t h e t r ansduce r s i n s t a l l e d on t h e loop and wi th t h e loop f u l l of mercury o r NaK t o e l i m i n a t e e r r o r s caused by s t a t i c head v a r i a t i o n s . Other s l a c k diaphragm p re s su re t r ansduce r s a r e used t o measure abso lu t e p re s su re a t t h e NaK flow i n l e t ( P l ) , mercury pump e x i t (P6), b o i l e r i n l e t (PS), b o i l e r o u t l e t (P3) and condenser i n l e t (P5). A l l of t h e s e t r ansduce r s funct ioned t h e same a s t h e d i f f e r e n t i a l t r ansduce r s descr ibed above except t h a t t hey provided a 4 - 20 ma dc s i g n a l d i r e c t l y . A l l except t h e gauge (P5) on t h e condenser were c a l i b r a t e d p r i o r t o i n s t a l l a t i o n i n t o t h e loop system a t ope ra t i ng temperatures . Liquid l e v e l gauges were a l s o provided i n t h e NaK dump tank , mercury dump t ank , NaK s t andp ipes and t h e mercury condenser t o permit accu ra t e c o n t r o l of inventory du r ing opera t ion . ,%utomatic loop s a f e t y c o n t r o l c i r c u i t s were i n s t a l l e d on t h e c o n t r o l console . These provide automatic dump of t h e mercury loop i f p ressure l i m i t s a t t h e b o i l e r i n l e t o r t h e condenser a r e exceeded and automatic shutdown of t h e gas - f i r ed furnace and a l l mercury loop h e a t e r s i f t h e NaK flow drops below a prcsci l e v e l or i f t h e gas h e a t e r t ube temperature excecds a p r e s e t l i m i t . Automatic c o n t r o l c i r c u i t s were a l s o provided f o r t empera tures of t h e NaK dump t ank h e a t e r , mercury vapor l i n e h e a t e r , mercury p rehea t e r and t h e gas h e a i e r e x i t . Addi t iona l warning dev ices a r e connected t o an annuncia tor c i r c u i t l oca t ed on t h e c o n t r o l console t o a l e r t loop opera lors o f t h e fol lowing condi t ions : 1. Gas-f i r e d furnace overtemperature 2. S t a t i c NaK expansion t ank overpressure 3. Mercury b o i l e r i n l e t overpressure 4. Mercury condenser overpressure A s shown i n Figure (14) , a d d i t i o n a l p re s su re t r ansduce r s P12, P13, P14, and P15 a r e used i n t h e argon and vacuum pip ing systems f o r loop ope r a t ion requirements . These a r e a l l convent ional penumat i c t r a n s m i t t e r s with a l l welded bourdon tube sens ing elements. A d e t a i l e d d e s c r i p t i o n of t h e thermocouples i n s t a l l e d i n t h e t e s t loop i s given i n Table ( I ) l i s t e d i n Appendix (11). B. B o i l e r I n s t rumentat i o n A s shown schemat ica l ly i n Figure (18), s i x t y 0.125-inch OD themo- couples were i n s t a l l e d on t h e s h e l l s u r f a c e of t h e t e s t b o i l e r i n 17 c i r c u m f e r e n t i a l r i ngs a t 12-inch i n t e r v a l s a long t h e e n t i r e boi l e s length. A 11 she 11 thermocouple junc t ions were formed by s u r f a c e spo t welding. I t was found t h a t a small t a b of 0.002-inch t h i c k tantalum f o i l tacked over t h e exposed junc t ions would reduce t h e frequency of thermocouple f a i l u r e caused by t h e thermocouple lead l i f t i n g o f f t h e s h e l l sur face . Seven 0.04-inch OD sheathed chromel-alumel thermocouples wi th capped and nongrounded junc t ions were i n s t a l l e d i n t h e centerbody of t h e h e l i c a l i n s e r t , wi th junc t ions loca ted a t evenly space increments along t h e a c t i v e b o i l i n g region of approxirnatcly t h e f i r s t 12 f e e t of t h e t e s t Output Signal , emf, mv Figure 17. Cal ibra t ions of Pressure Transducers. boi l e r . These i n s e r t thermocouples were then brought out through aUConax'hulti-hole packing gland s o t h a t t h e i n s i d e of t h e center- hocly tube could be kept under pressurized argon atmosphere during operat ion. This arrangement ensures t h a t argon w i l l leak i n t o t h e mercury loop i n t h e event of any leak along t h e centerbody tube, r a t h e r than mercury leaking i n t o t h e atmosphere. Two 0.062-inch OD Cr-A1 sheathed thermocouples wi th capped and nongrounded junct ions were posi t ioned i n each of the four themocouple wel ls located a t t h e i n l e t and o u t l e t of both t h e NaK and mercury flows. Two 0.125-inch OD Cr-A1 sheathed thermocouples 180 degrees apa r t were surface spot welded i n t h e i n l e t and o u t l e t of both the primary and secondary flow pipes t o measure t h e i n l e t and o u t l e t pipe sk in temperatures. Two s l ack diaphragm Taylor absolute pressure t ransducers located a t the mercury i n l e t and e x i t were employed t o measure t h e t o t a l pressure drop across the t e s t b o i l e r . Figure (13) shows-the mercury o u t l e t pressure transducer a s i n s t a l l e d i n t h e t e s t sec t ion . The d i f f e r e n t i a l pressure gauge used t o measure NaK flow pressure drop through t h e s h e l l s i d e passages was a l s o a s l ack diaphargmunit . A d e t a i l e d desc r ip t ion of the s h e l l and i n s e r t thermocouples used i n t h e t e s t sec t ion i s given i n Tables (2) and ( 3 ) and is l i s t e d i n Appendiq 11. C. Loop Checkout and Operat ing Procedures One of t h e problems of g r e a t e s t concern was t h a t of p a r t i c u l a t e ma t t e r which might be p re sen t i n t h e loop p r i o r t o s t a r t u p . Every p recau t ion was gaken t o e l i m i n a t e t h i s p o s s i b i l i t y , however, it must be apprec ia ted t h a t d e s p i t e a l l c a r e t h e p o s s i b i l i t y s t i l l e x i s t e d . Consequently, p r i o r t o s t a r t u p t h e loop was f lushed wi th mercury i n both d i r e c t i o n s and t h e mercury was r e tu rned t o t h e dump tank . The d i p l e g i n t o t h e dump t ank was t hen blown down wi th p re s su r i zed argon t o make s u r e t h a t any p a r t i c u l a t e m a t e r i a l f l o a t i n g on t h e s u r f a c e was blo&n i n t o t h e dump tank . The p r o b a b i l i t y of such contaminants being re turned t o t h e loop on r e f i l l i n g w i l l t h u s be extremely smal1,s ince t h i s m a t e r i a l would t h e n be f l o a t i n g on t h e s u r f a c e of t h e mercury i n t h e dump tank . I n a d d i t i o n t o t h i s f l u s h i n g procedure, a f i l t e r having a nominal pore s i z e of 17 microns was placed i n t h e loop j u s t upstream of t h e l i q u i d t h r o t t l e va lve where t h e l i q u i d mercury e n t e r s t h e b o i l e r . The f i l t e r should prevent p a r t i c l e s from reaching and plugging t h e 0.060-inch o r i f i c e s of t h e i n l e t of each b o i l e r t ube . Oxide and impuri ty c o n t r o l i n t h e NaK primary loop was achieved by a combination of ho t - f l u sh ing and hot - t rapping . Upon s t a r t u p , t h e NaK loop was hot-f lushed a t 800 OF by s e v e r a l charges of NaK flow, which was d i scarded . Oxygen conten t dur ing ope ra t i on was c o n t r o l l e d by hot - t rapping t h e l i q u i d NaK a t 1200°F wi th zirconium g e t t e r i n g g r i d Located i n t h e NaK dump tank. The s t a t i c NaK l a y e r system which ac ted a s a b a r r i e r between t h e primary NaK flow and t h e secondary mercury flow was f i l l e d wi th 38 793 grams of reactor-grade NaK. The s t a t i c NaK system was vacuum out- gassed f o r 36 hours a t temperatures up t o 700°F be fo re f i l l i n g . During t h i s time, t h e mercury system was a l s o under vacuum t o p r o t e c t t h e tantalum t e s t b o i l e r from contamination. F i l l i n g ope ra t ion proceeded from a cond i t i on i n which both t h e primary and secondary loops were evacuated wi th t h e f a c i l i t y mechanical 0 vacuum pump and were baked o u t t o approximately -. 400 F by means of a u x i l i a r y e l e c t r i c a l hea t ing wire . The mercury and NaK dump t anks were a l s o heated up t o approximately 400°F with a u x i l i a r y hea t ing equipment. Then by p re s su r i z ing .the NaK dump t ank with argon, NaK was forced t o f i l l t h e evacuated loop. Level s enso r s loca ted i n t h e NaK loop st andpipes i nd ica t ed when t h e proper inventory was obtained. I n t h e meantime, t h e NaK EM pump was a c t i v a t e d and, subsequently, t h e gas- f i r e d furnace was turned t o i t s minimum hea t l e v e l . The secondary mercury loop was f i l l e d i n a s i m i l a r manner. The l i q u i d mercury was forced t o f i ' l l t h e evacuated mercury loop by p re s su r i z ing argon i n t h e mercury dump. When t h e d e s i r e d mercury inventory was a t t a i n e d , t h e mercury EM pump was then a c t i v a t e d and, i n t h e meantime, forced coo l ing a i r flow i n t o t h e condenser was s t a r t e d . Bo i l i ng was i n i t i a t e d a s t h e mercury was pumped i n t o t h e NaK heated b o i l e r s e c t i o n where t h e s h e l l s i d e NaK flow had a l r eady been heated t o an e leva ted temperature. Before f i l l i n g with mercury, a s e r i e s of thermocouple c a l i b r a t i o n s and hea t l o s s . runs wereconducted. The purpose of t h e s e runs was t o check each b o i l e r s h e l l thermocouple aga ins t t h e c a l i b r a t e d w e l l thermocouple i n t h e NaK stream and t o o b t a i n t h e heat l o s s f o r t h e test s e c t i o n . Th i s was accomplished by running t h e NaK loop a t both maximum and minimum I l o w r a t e s a t t h r e e d i f f e r e n t temperature l e v e l s wi th t h e mercury loop evacuated. The hea t l o s s runs a r e of v i t a l importance i n t h e heat t r a n s f e r a n a l y s i s t o determine accu ra t e ly t h e exac t amount of heat t r a n s f e r r e d t o t h e tantalum b o i l e r t ube . A t echnique descr ibed i n d e t a i l i n Reference (6 ), employing two NaK flow r a t e s a t s e v e r a l temperature l e v e l s with t h e mercury loop evacuated, was used t o determine t h e t e s t s e c t i o n hea t l o s se s . These hea t l o s s e s , s o determined, a r e presented i n F igure (19) and range from 6.26 t o 7.4 KW depending upon t h e temperature l e v e l of t h e t e s t s e c t i o n . The s h e l l thermocouple c o r r e c t i o n s ob ta ined i n t h i s manner were g e n e r a l l y about 10°F and a r e + bel ieved accu ra t e t o approximately - 2OF. The r e f e r ence wel l thermocouples + i n t h e NaK stream a r e judged t o have an accuracy w i t h i n - 1°F on an abso lu t e bas i s . I V . EXPERIMENTAL RESULTS AND DISCUSSIONS Before present ing average and l o c a l h e a t t r a n s f e r and p re s su re drop r e s u l t s , a conceptual model of t h e l o c a l b o i l i n g cond i t i ons as shown i n F igure (20) which i s thought t o occur i n t h e mercury once- through b o i l e r w i l l be b r i e f l y descr ibed i n o rde r t o d e f i n e some of t he terminology employed i n t h i s r e p o r t . In t h e once-through b o i l i n g process a s descr ibed i n Reference ( 6 ) . t he mercury bulk temperature, subcooled a t t h e b o i l e r i n l e t , i s increased by s i n g l e phase l i q u i d hea t t r a n s f e r i n t h e "subcooled hea t ing region" u n t i l b o i l i n g is i n i t i a t e d . The poin t of b o i l i n g incep t ion marks t h e beginning of t h e "nucleate b o i l i n g region1' which i s mainly cha rac t e r i zed by r e l a t i v e l y high hea t t r a n s f e r performance. I n t h i s reg ion , it i s bel ieved t h a t t h e w a l l is completely wetted by mercury a t t h e e l eva t ed s a t u r a t i o n temperatures of t h e present t e s t . A s t h e q u a l i t y is increased , it is thought t h a t i n t h i s nuc lea t e reg ion pa r t of t h e l i q u i d flows a s a continuous l i q u i d f i l m on t h e b o i l e r wal l and t h e remainder is en t r a ined i n t h e vapor core. Two hea t t r a n s f e r models a r e f r equen t ly used t o desc r ibe t h i s nuc lea te b o i l i n g reg ion a t high q u a l i t i e s . The " f i lm evaporat ionf ' model, which is cha rac t e r i zed by t o t a l l y suppressed bubble formation has t h e vapor generated by evapora t ion a t t h e pos tu l a t ed i n t e r f a c e . The "bubble nuc lea t ion" model neg lec t s l iquid-vapor i n t e r f a c e vapor gene ra t ion and assumes hea t t r a n s p o r t through t h e formation of vapor bubbles from n u c l e a t i o n s i t e s a t t h e b o i l e r tube wal l . Th i s nuc l ea t e b o i l i n g r eg ion p e r s i s t s with i nc reas ing q u a l i t y u n t i l t h e onset of t h e c r i t i c a l heat f l u x cond i t i on , a t which poin t t h e hea t t r a n s f e r performance begins t o d e t e r i o r a t e due t o breakdown of t h e cont inuous l i q u i d f i l m along t h e b o i l e r wal l . Thus, t h e c r i t i c a l hea t f l u x cond i t i on marks t h e onset of t h e lower hea t t r a n s f e r per f ormancel ' t ransi t i o n b o i l i n g regiont: The t r a n s i t i o n b o i l i n g reg ion i s cha rac t e r i zed by a lower hea t t r a n s f e r c o e f f i c i e n t thought t o be caused by t h e hea t t r a n s f e r s u r f a c e be ing p a r t i a l l y wetted by l i q u i d pa tches o r d r o p l e t s and p a r t i a l l y exposed t o d ry vapor. A s t h e q u a l i t y i s f u r t h e r increased , t h e q u a n t i t y of l i q u i d pa tches o r d r o p l e t s wet t ing t h e wa l l decrease and t h e mean w a l l temperature i n c r e a s e s r a p i d l y u n t i l t h e so-ca l led "sphero ida l stateq' o r " f i lm b o i l i n g region" i s reached. This reg ion i s cha rac t e r i zed by t h e wal l being blanketed wi th a cont inuous l a y e r of l o c a l l y super- heated vapor. The f i lm b o i l i n g reg ion ex tends t o t h e po in t a t which bulk superhea t ing commences, and t h e subsequent "superheated vapor regionv' marks t h e end of t h e once-through b o i l i n g process . I n t h e f i lm b o i l i n g reg ion , hea t i s t r a n s f e r r e d by a combination of conduction, convect ion and r a d i a t i o n through t h e vapor l a y e r , and t h e hea t f l u x e s expected a t reasonable temperature l e v e l s a r e sub- s t a n t i a l l y lower t han those i n t h e nuc l ea t e b o i l i n g region. For a h o r i z o n t a l su r f ace , t h e f i l m b o i l i n g hea t t r a n s f e r c o r r e l a t i o n s have been worked out by s e v e r a l i n v e s t i g a t o r s . Chang( "and Zuber ( 91 a r e among t h e e a r l i e s t workers. I n 1961,Berensen ( lo ) modified and extended t h e hydrodynamic approach used by Zuber and Chang and obta ined r e l a t i o n s h i p s f o r both t h e hea t t r a n s f e r c o e f f i c i e n t and t empera ture d i f f e r e n c e a t t h e minimum hea t f l u x cond i t i on f o r pool bo i l i ng . The r e s u l t s a r e given a s fol lows: The parameter (AT) i s t h e minimum temperature d i f f e r e n c e between m i n b o i l e r wa l l and mercury requi red f o r t h e e x i s t e n c e of f i l m b o i l i n g . Equation (1) was used t o p r e d i c t t h e t ube wall-to-mercury temperature d i f f e r e n c e r equ i r ed f o r mercury having f i l m b o i l i n g . I t i s found t h a t T . i s g e n e r a l l y i n t h e o r d e r of 250°F under t h e p re sen t ope ra t i ng m i n cond i t i ons . One should n o t i c e t h a t Equat ion (1) i s p r i m a r i l y der ived f o r pool b o i l i n g and it might not be u s e f u l t o p r e d i c t (AT) . f o r ml n onse t of f i l m b o i l i n g under fo rced convect ion condi t ions . However, due t o l ack of in format ion f o r forced-convect ive f i l m b o i l i n g f o r mercury, it i s be l ieved t h a t Equat ion (1) w i l l be adequate t o provide a t l e a s t an o rde r of magnitude ana lys i s . For tuna te ly , r ecen t f i l m b o i l i n g d a t a f o r mercury from Reference (11) showed t h a t t h e mean temperature d i f f e r e n c e f o r f i l m b o i l i n g i s i n t h e range of 180 - 2T0°F under t h e same tempera ture l e v e l a s t h e p re sen t experiment employedl. These sources suppor t a conclusion t h a t t h e f i l m b o i l i n g might not have occurred i n t h e p re sen t two-f luid mercury b o i l e r , s i n c e no tube wall-to-mercury tempera ture d i f f e r e n c e s a s l a r g e a s 200°F were experienced i n t h e p re sen t experiment. A. Overa l l and Average Tes t Resu l t s The d i g i t a l d a t a system, a s p rev ious ly descr ibed , rece ived and recorded raw test d a t a i n t h e form of a m i l l i v o l t s i g n a l . A computer program was w r i t t e n t o convert t h e s e q u a n t i t i e s t o u s e f u l engineer ing u n i t s which would permit d i r e c t use of t h e convent iona l equa t ions f o r performance a n a l y s i s . The computations used i n t h e computer program make use of a l l a v a i l a b l e c a l i b r a t i o n f a c t o r s f o r th(~rmocouples , p ressure gauges, and flowmeters. A l l phys ica l q u a n t i t i e s f o r mercury l i q u i d and vapor used i n t h e da t a r educ t ion a r c obtained from Reference (12) . Appendix (111) shows t h e p r in tou t of most of t h e o v e r a l l t e s t r e s u l t s f o r t h e presen t 16 t e s t runs , i nc lud ing such f a c t o r s a s power l e v e l s , ope ra t i ng temperature bands, degree of superhea t ing , flow r a t e s , and t o t a l p r e s su re drops. Indeed, t h e s e o v e r a l l r e s u l t s have i l l u s t r a t e d t h e gross performance f o r b o i l i n g of mercury i n s i d e a tan ta lum tube b o i l e r . F igures (21) t o (36) show t h e temperature p r o f i l e s f o r NaK and mercury flows i n t h e test s e c t i o n f o r a l l 16 t e s t runs. Ex i t superhea ts of t h e mercury vapor a r e a l s o i l l u s t r a t e d i n t h c s e f i g u r e s . Overa l l r e s u l t s of t h e ope ra t i ng temperature band f o r primary and secondary flows, a s shown i n thcse f i g u r e s , a r e q u i t e u s e f u l i n t h a t they i n d i c a t e t h e f e a s i b i l i t y of compact once-through b o i l e r s and permit t h e design of mul t ip le - tube b o i l e r s with enhanced thermal power l e v e l s by d i r e c t scale-up. For example, t h e des ign of a compact once-through 19-tube boi ler(13)which was proposed f o r t h e SNAP-8 power conversion system was based on t h e o v e r a l l r e s u l t of a 1 /19 th scale t e s t run (F igure 35). This compact 19-tube b o i l e r w i t h only 6.68 feet long w i l l meet t h e SNAP-8 power conversion system requi re - ments, i . e . , a power l e v e l of 500 KW and a s p e c i f i e d e x i t pinch p o i n t temperature d i f f e r e n c e of 3 0 ' ~ . Following t h e computat i o n a l procedures given i n Appendix ( I ) , t h e average values of z, h, and q" f o r each i n d i v i d u a l hea t t r a n s f e r reg ion were c a l c u l a t e d , and t h e r e s u l t s f o r a l l 16 test runs, a r e l i s t e d i n Table ( 4 ) . The ranges of t h e average values are as follows: The l eng th f o r each i n d i v i d u a l h e a t t r a n s f e r r eg ion except t h e subcooled hea t ing r eg ion was ca l cu la t ed fol lowing t h e procedures presented i n Appendix (I). The l e n g t h of t h e subcooled l i q u i d hea t ing reg ion , LSC. was obtained d i r e c t l y from t h e mercury temperature p r o f i l e determined from t h e i n s e r t thermocouples. A s shown i n Figures (21) t o ( 3 6 ) , t h e subcooled hea t ing reg ion is gene ra l ly s h o r t ( l e s s than 7% of t h e t o t a l b o i l e r l eng th ) f o r a l l t e s t runs. The values of o v e r a l l p r e s su re drops recorded from t h e two Taylor gauges a r e l i s t e d i n Table (5) and a r e p l o t t e d i n F igures (21) t o (36) . The p re s su re drop p r o f i l e s in Figures (21) t o (36) were obtained w i t h t he s a t u r a t i o n curves from t h e s a t u r a t i o n temperature of mercury measured wi th i n s e r t thermocouples. The d o t t e d l i n e s i n t h e s e f i g u r e s a r e e s t ima te s f o r mercury vapor i n t h e superheated reg ion . Following t h e c a l c u l a t i o n a l procedures desc r ibed i n Appendix ( I ) f o r (AP)misC., (@) elev, ) (Ap) SpL) and (Ap) S p V ~ t h e two p re s su re drops were ob ta ined , and t h e s e r e s u l t s a r e given i n Table ( 5 ) . The va lues of (AP) obtained TP from s a t u r a t i o n curves by i n s e r t thermocouples read ings a r e a l s o l i s t e d i n Table ( 6 ) f o r comparison t o t hose d a t a obtained from Equation (A38). A s shown i n F igure (37 ) , t h e comparison i n d i c a t e s t h a t the va lues of (&) TP obtained from two d i f f e r e n t sources g e n e r a l l y agreed + 16%. I n F igure (38) - t h e f r i c t i o n a l and momentum components of (AP) a r e p l o t t e d along wi th TP t h e value of (AP) obtained from i n s e r t thermocouples. The momentum TP component (Ap) i s c a l c u l a t e d from Equation (A57). TPrn The va lues of (Ap) obtained from i n s e r t thermocouples were TP used i n Equation (A38), and t h e mercury vapor p re s su re drop, (AP) SPV Was obtained. By us ing Equation (A39), t h e f r i c t i o n a l term, was t hen obtained by s u b t r a c t i n g t h e momentum t e r m , L I P ~ ~ ~ m . Again a comparison between t h e s e experimentally-determined (@) SPVf and t h e va lues c a l c u l a t e d by Equations (A46) and (A53) was made and t h e r e s u l t i s presented i n Figure (39) . The comparison i n d i c a t e s t h a t t h e s e two (Ap) SPVf g e n e r a l l y agree w i t h i n + 22% and t h e p red i c t ed va lues a r e g e n e r a l l y sma l l e r t h a n - those obtained from t h e t e s t r e s u l t s . The momentum p a r t s c a l c u l a t e d by Equation (A41) a r e found t o be a sma l l f r a c t i o n of t h e t o t a l mercury vapor p re s su re drop ( l e s s t h a n lo%) , and t h e i r va lues a r e p l o t t e d i n F igure (40) t o g e t h e r wi th (Ap) (Ap) sPV' and t o t a l p r e s su re drop ac ros s t h e b o i l e r TP , ( /P)T. The r a t h e r l a r g e - + 22% discrepancy i n (AP) SPVf between t e s t e d and p red i c t ed values i s a t t r i b u t e d mainly t o t h e u n c e r t a i n t y of u s ing t h e f r i c t i o n a l p ressure m u l t i p l i e r K e s p e c i a l l y i n t h e wi re c o i l reg ion . The use of Equation P' (A52) f o r p r e d i c t i n g K va lues i n t h e h e l i c a l vane i n s e r t r eg ion has r e c e n t l y P been v e r i f i e d by a s e r i e s of water and a i r tests . (29) I n o t h e r words, t h e e m p i r i c a l K c u r v e ob ta ined f o r a wire c o i l i n r e f e r e n c e ( 2 9 ) i s b~lic.:*cxl 1" t n be c o r r e c t wi r i i ln 5 22% when u s i n g i t t o p r e d i c t f r i c t i o n a l pressure l o s s of mercury vapor fl-owing through t h e wire c o i l i n s e r t rcgior l , It was a l s o found t h a t a r a t h e r l a rg i? f r a c t i o n of p r e s s u r e d r c ~ occur red i n t h e w i r e c o i l i n s e r t r e g i o n as c o q a r e d w i t h o t h e r portCo:.s a long t h e b o i l e r . T h i s i s c l e a r l y sho~+-i'. I n F i g u r e s (21) t o (363, A- w i r e c o i l i n s e r t a t t h e end of t h e boller ::he superhea ted reg ion) b a s used t o t u r b u l a t e t h e l i q u i d mercury v h i c h ~ . i g h t remain on t h e c e n t e r - body a t t h e h e l i c a l vane i n s e r t end and zz enhance t h e superhea ted T d z s o r h e a t t r a n s f e r c o e f f i c i e n t . A s t h e p r e s s - r e Zrop d a t a shows, t h e 7eazlt.y f o r t h i s p e r f o r n a n c e was an i n c r e a s e of ;-3;,~in g power f o r t h e workl=g f l u i d . B. Local Tes t Resu l t s The o v e r a l l and average hea t t r a n s f e r and p re s su re drop d a t a presented prev ious ly a r e only u s e f u l f o r des ign s t u d i e s a t cond i t i ons near those a t which t h e t e s t d a t a were obtained. These d a t a were gene ra l l y obtained by averaging over one o r more hea t t r a n s f e r reg ions of t h e once-through b o i l i n g process and thus can not be ex t r apo la t ed w i th confidence f o r gene ra l des ign purposes beyond t h e range of t h e presen t t e s t condi t ions . For t h e s e reasons, l o c a l mercury b o i l i n g heat t r a n s f e r and pressure drop d a t a were needed, and consequently a cons iderab le e f f o r t was devoted t o t h e c a l c u l a t i o n and c o r r e l a t i o n of t h e l o c a l r e s u l t s . These l o c a l r e s u l t s not only enhanced understanding of t h e once-through b o i l i n g mercury i n a s i n g l e tube wi th i n s e r t s , but a l s o provided a gene ra l means f o r e x t r a p o l a t i n g d a t a o u t s i d e t h e range of t e s t cond i t i ons . 1. Heat Flux D i s t r i b u t i o n , Q u a l i t y and Overa l l Heat T rans fe r C o e f f i c i e n t Values of t h e l o c a l hea t f l u x d i s t r i b u t i o n were obtained by d i f f e r e n t i a t i o n of t h e smooth temperature curves a t one-foot i n t e r v a l s , us ing Equation (A64) f o r t h e reg ions up t o t h e t r a n s i t i o n b o i l i n g and Equation (A66) f o r t h e reg ion of superheated vapor. The r e s u l t s a r e presented i n F igures (41) t o (56) . The values of q" obtained i n t h i s manner were not s u f f i c i e n t t o g ive t h e e n t i r e a x i a l hea t f l u x p r o f i l e , s i n c e near t h e end of t h e a c t i v e heat t r a n s f e r l ength t h e l o c a l d i f f e r e n t i a - t i o n of t h e mercury temperature p r o f i l e s could not be made wi th accuracy. However, wi th t h e known values of t h e average hea t f l u x i n t h e superheated reg ion , t h e shape of l o c a l hea t f l u x d i s t r i b u t i o n s could be es t imated i n t h e end reg ion . The es t imated p r o f i l e s a r e shown, i n F igures (41) t o ( ~ 3 6 ) ~ Also shown i n these f i g u r e s , t h e p o i n t s I B , C , and I S i n d i c a t e the approximate l o c a t i o n of b o i l i n g incep t ion , c r i t i c a l h e a t flux cond i t i on , and i n i t i a t i o n of superhea t ing f o r each t e s t run. I t can be seen from t h e s e f i g u r e s t h a t , i f t h e "entrance e f f ec t ' ' is excluded before and a f t e r t h e s e po in t s , t he l o c a l hea t f l u x i n t h e ind iv idua l heat t r a n s f e r reg ions inc reases o r decreases almost l i n e a r l y . This c f f c c t can be more c l e a r l y seen from those runs with long b o i l i n g s e c t i o n s i n F igures (52 ) . (53 ) ) and (54) . Local q u a l i t i e s ca l cu la t ed from Equation (A70) a r e p l o t t e d i n F igures (21) t o (36). The r e s u l t s show t h a t t h e convent ional assumption of l i n e a r v a r i a t i o n i n q u a l i t y versus length is not always c o r r e c t , e s p e c a i l l y f o r t h e low and high q u a l i t y reg ions and is acceptab le only i n t h e in te rmedia te q u a l i t y range, i . e . , 0 .2 t o 0.8. Local values of t h e o v e r a l l hea t t r a n s f e r c o e f f i c i e n t , U, were ca l cu la t ed from Equation (A73), and t h e r e s u l t s a r e presented i n Figures (41) t o (56) . Due t o t h e l a c k of r e l i a b l e q" values near t h e end of t h e t e s t s e c t i o n , t h e d i s t r i b u t i o n of U is extended by dot ted l i n e s t o t h e known value a t t h e b o i l e r e x i t . A t t h e b o i l e r e x i t , t h e superheated vapor heat t r a n s f e r c o e f f i c i e n t was est imated by t h e convent ional Di t tus -Boel te r equat ion, modified f o r h e l i c a l flow a s follows: where t h e s u b s c r i p t b denotes t h e f l u i d p r o p e r t i e s eva lua ted a t f l u i d bulk temperature . The l o c a l va lues of U were then used t o c a l c u l a t e t h e l o c a l mercury hea t t r a n s f e r c o e f f i c i e n t s from Equation (A74). A l l t h e l o c a l h r e s u l t s l-i g were t abu la t ed f o r each ind iv idua l hea t t r a n s f e r r eg ion (Tables 7 , 8 , 10, 11) I n t h e fol lowing s e c t i o n , a more d e t a i l e d d i s cus s ion of t h e s e h va lues Hg is given. 2. Local Heat T rans fe r C o e f f i c i e n t s and Cor re l a t i ons Subcooled Liquid Region, h~ Resu l t s of subcooled l i q u i d hea t t r a n s f e r c o e f f i c i e n t s a r e summarized i n Table (7 ) . F igure (57) shows a comparison of t h e presen t d a t a t o t h e convent ional p r e d i c t i o n s by Lyon ( I 4 ) and Lubarsky and Kaufman (15) . When w r i t t e n i n terms of h e l i c a l flow parameters , t h e s e two equa t ions a r e b ~ D e -= 7 + 0.025 (Lyon' s Equation) k~ 0 .4 h D = 0.625 G ~ D e C PL k k (Lubarsky - Kaufman Equation) (4) L L A s shown i n t h e f i g u r e , t h e mercury d a t a f i t w e l l wi th t h e Lubarsky-Kaufman equa t ion when t h e P e c l e t number, * ~ e L was l e s s t han t 250; when t h e Pec l e t number was l a r g e r t han 250, t h e d a t a were c l o s e t o t h e p r e d i c t i o n of yon's equa t ion . Nucleate Bo i l i ng Region, hm Six ty - s ix l o c a l nuc l ea t e b o i l i n g d a t a p o i n t s were obtained by t h e procedure d i scussed i n Appendix (I) and were t a b u l a t e d a able 8 ) . F igure (58) p r e s e n t s t h e c a l c u l a t e d h e a t t r a n s f e r c o e f f i c i e n t s i n t h e nuc l ea t e b o i l i n g region,hm, versus t h e l o c a l hea t f l u x , q", wi th mass v e l o c i t y , G , and s a t u r a t i o n temperature , T used a s parameters. Hg s a t ' From t h e s e curves, it i s c l e a r t h a t t h e mass v e l o c i t y and s a t u r a t i o n temperature ( o r p ressure) both have an e f f e c t i n i nc reas ing t h e nuc l ea t e b o i l i n g hea t t r a n s f e r c o e f f i c i e n t . This can be s een more c l e a r l y by comparing curves (5) and (7) f o r t h e mass v e l o c i t y e f f e c t , and curves (7 ) and (8) f o r t h e b o i l i n g temperature e f f e c t . F igure (59) p re sen t s q" versus AT, where AT, t h e temperature d i f f e r e n c e between t h e w a l l temperature and f l u i d s a t u r a t i o n temperature , was obtained by t h e fo l lowing equat ion: The approximate l i q u i d met a 1 forced convect ion hea t t r a n s f e r equa t ion (Nu = 7) and a v a i l a b l e mercury pool b o i l i n g d a t a a r e a l s o presented i n F igure (59) f o r comparison. The pool b o i l i n g curve and d a t a p o i n t 0 a t T = 950 F were c i t e d from t h e t e s t r epo r t ed by Atomics I n t e r - S a t n a t i o n a l ( I 6 ) . Another s e t of pool b o i l i n g curves f o r e l eva t ed s a t u r a t i o n temperature l e v e l s were ob t a ined by extending t h e Russian pool b o i l i n g equat ion(17) a t 14.7 p s i a o r 680'~. The pool b o i l i n g equa t ion ( 1 7 ) reads : The C iche l l i -Bon i l l a c o r r e l a t i o n (18) . 5 was employed t o ob ta in : t I 11 (AT > = q c - PB AT C where t h e c r i t i c a l hea t f l u x and AT were obtained from Zuber 's Equation ( 19) C f o r s e l e c t e d s a t u r a t i o n temperatures . To examine t h e dependence of h on hea t f l u x and q u a l i t y , h was NB NB p l o t t e d versus q" f o r s e l e c t e d d a t a ponts having approximately t h e same q u a l i t i e s , and verus x f o r s e l e c t e d d a t a po in t s having approximately t h e same hea t f l uxes . F igure (60) shows t h a t h i nc reases w i t h i nc reas ing NB heat f l u x , whereas Figure (61) shows t h a t h is v i r t u a l l y independent NB of t h e l o c a l q u a l i t i e s f o r s e l e c t e d cons t an t hea t f l u x l e v e l s . Moreover, t h e i nc rease i n h w i t h mass v e l o c i t y a s shown i n t h e s e f i g u r e s can be m a t t r i b u t e d t o a hea t f l u x e f f e c t s i n c e , i n t h e two-f luid hea t exchanger, hea t f l u x i nc reases i n proporat i o n t o mass v e l o c i t y a t cons tan t q u a l i t y . To a s c e r t a i n i f any d e t e c t a b l e t r end e x i s t e d i n t h e d a t a of h a s a func t ion of r a d i a l a c c e l e r a t i o n developed by t h e h e l i c a l N B * i n s e r t , Figure (62) was cons t ruc ted by p l o t t i n g h versus ( 1 + NB Pk) where A denotes t h e r a d i a l a c c e l e r a t i o n of t h e annular l i q u i d a t t h e R ( 6 ) a wal l s u r f a c e and i s g iven by t h e fol lowing express ion . As shown i n Figure (62 ) ) t h e e f f e c t of l o c a l r a d i a l a c c e l e r a t i o n A is R an inc rease i n t h e nuc lea t e b o i l i n g hea t t r a n s f e r c o e f f i c i e n t a t t h e Pow 2 a c c e l e r a t i o n f i e l d o r t h e low q u a l i t y region, x < O . 10 ( s i n c e A w x ) " R I t can be seen from t h i s f i g u r e t h a t t h e degree of i nc reas ing h by NB r a d i a l a c c e l e r a t i o n i s g radua l ly decreased a s t h e q u a l i t y is increased . For x > 0 . 3 5 , t h e r e is almost no e f f e c t on h NB by AT{. Somewhat s i m i l a r l y , t h e pool b o i l i n g da t a of Merte and Clark(20)showed t h e e f f e c t of a c c e l e r a t i o n f i e l d on nuc lea te pool b o i l i n g of water . The i r r e s u l t s showed an inc rease i n h wi th inc reas ing a c c e l e r a t i o n f i e l d s a t low heat f l u x e s , NB and a decrease i n h wi th inc reas ing a c c e l e r a t i o n f i e l d s a t high heat NB f l uxes . However, i n t h e present experiments t h e r eve r se e f f e c t a t h igher q u a l i t i e s was not observed, i . e . , hNB d id not decrease a s t h e a c c e l e r a t i o n f i e l d increased . The evidence i n F igures (60) and (61) t h a t h was s t r o n g l y inf luenced NB by heat f l u x and much l e s s inf luenced by q u a l i t y , l ed t o a f u r t h e r i nves t iga - t i o n of t h e heat t r a n s f e r mechanism of wet ted mercury b o i l i n g i n s i d e a tantalum tube . As noted i n Reference (21) , a b o i l i n g mechanism termed a s " f i lm evapora t ionw, which is based upon conduction through t h e t h i n l i q u i d f i l m wi th subsequent evapora t ion from t h e f i lm s u r f a c e , p r e d i c t s t h e h to m be markedly inf luenced by q u a l i t y and t o decrease with inc reas ing temperature leve 1s. References (2 1) and (22)note another b o i l i n g mechanism, "bubble nucleat ion", i n which h inc reases wi th both heat f l u x and temperature NB l e v e l . Hence, F igures (60) and (61) suggest t h a t t h e r e may be bubble nuc lea t ion f o r mercury i n forced convect ion bulk bo i l i ng . Below i s a more d e t a i l e d a n a l y t i c a l s tudy of t h e s e two hea t t r a n s f e r models. The a n a l y t i c a l flow p a t t e r n is assumed t o c o n s i s t of a t h i n , continuous, and concentric l aye r of mercury l i q u i d on t h e wa l l , wi th t h e remaining l i q u i d en t ra ined 56 i n t h e vapor core and t r a v e l i n g wi th t h e vapor. Film Evaporation Model - The hea t t r a n s f e r mechanism i s assumed t o be conduction from t h e w a l l t o t h e l iquid-vapor i n t e r f a c e through t h e t h i n l aye r of l i q u i d on t h e wa l l . The vapor is then generated by evapora t ion a t e x i s t i n g l iquid-vapor i n t e r f a c e s . Evaporation of t h e en t r a ined l i q u i d d r o p l e t s i s neglec ted , and t h e i n t e r f a c e is assumed t o be a t t h e l o c a l s a t u r a t i o n temperature . By u t i l i z i n g t h e hea t conduction equa t ion through t h e annulus l i q u i d l a y e r , t h e fol lowing express ion was obtained : where D denotes t h e diameter of t h e vapor core . v Equation ( l o ) , r e w r i t t e n i n terms of t h e average void f r a c t i o n and mass f r a c t i o n of t h e en t r a ined l i q u i d is a s follows: where I< = t h e s l i p r a t i o obtained from t h e c o n t i n u i t y equa t ion E = t h e f r a c t i o n of t h e en t r a ined l i q u i d flowing i n t h e vapor core . By assuming E = 0, Equation (11) reduced t o t h e fo l lowing simple express ion: The r e l a t i o n s h i p between q u a l i t y and void f r a c t i o n is given by t h e momentum exchange model from Reference ( 2 3 ) : By simply l e t t i n g t h e s l i p r a t i o equal one, t h e fol lowing'was obtained: P~ - ,= 1 (homogeneous model) Equations (14) and (15) a r e p l o t t e d i n Figure (63) f o r each s a t u r a t i o n temperature l eve l . The void f r a c t i o n a was t h e n obtained from Figure (63) by knowing t h e l o c a l q u a l i t y . Subsequently, t h e hea t t r a n s f e r c o e f f i c i e n t due t o f i l m evaporat i on a lone was c a l c u l a t e d from Equat ion ( 13) . To ob ta in agreement between t h e hea t t r a n s f e r c o e f f i c i e n t a t ze ro q u a l i t y and t h e f u l l y developed s ingle-phase l i q u i d value, an i n t e r p o l a t i o n equat ion was proposed t o o b t a i n t h e hea t t r a n s f e r c o e f f i c i e n t f o r t h e f i l m evapora t ion model i n t h e nuc lea te b o i l i n g region. This equat ion reads: with h ca l cu la t ed by Equation (3 ) . A s i nd i ca t ed by Equation (16) , h L TP approaches t h e s ingle-phase l i q u i d value a t low q u a l i t i e s and approaches t h e f i lm evaporat ion va lue a t high q u a l i t i e s . Bubble Nucleat ion Model - I n t h i s model, vapor i s assumed t o be generated by t h e formation of bubbles a t t h e wa l l of t h e tube, wi th subsequent growth and t r a n s p o r t of t h e s e bubbles through t h e l i q u i d 58 l aye r i n t o t h e vapor core. The heat t r a n s f e r c o e f f i c i e n t was obtained by a supe rpos i t i on of nuc lea te pool b o i l i n g and l i q u i d (24) forced convection a s recommended by Kutateladze , i . e . , : where h and h were evaluated by Equations (3) and (6 ) , r e spec t ive ly . L PB The ca l cu la t ed h from Equations (16) and (17) were used f o r com- TP h par i son with t h e present nuc lea t e b o i l i n g da ta . The r a t i o .of exp - h~~ i s shown versus vapor q u a l i t y and versus heat f l u x i n F igures (64) and ( 6 5 ) ) r e spec t ive ly . A s shown i n t h e s e f i g u r e s , t h e d a t a c o r r e l a t e d with t h e f i lm evaporat ion model a r e more s c a t t e r e d than those c o r r e l a t e d with .+ t h e bubble nuc lea t ion model. I n F igure ( 6 5 ) , t h e well-grouped da t a of t h e bubble nuc lea t i on model show t h e independence of vapor q u a l i t y . This s t r o n g l y i n d i c a t e s t h a t t h e mechanism of forced-convection b o i l i n g mercury might be t h e bubble nuc lea t ion i n t h e nuc lea t e b o i l i n g region. Furthermore, based upon t h e present r e s u l t s , a r a t h e r genera l ized c o r r e l a t i o n equat ion f o r mercury hea t t r a n s f e r c o e f f i c i e n t i n t h e nuc lea t e b o i l i n g r eg ion is: with h L' h~~ given i n Equations (3 ) and ( 6 ) , r e spec t ive ly . A c .or re la t ion of t h e nuc lea t e b o i l i n g hea t t r a n s f e r c o e f f i c i e n t s was attempted, based upon t h e fol lowing reasoning : 59 ( 1 ) The r e s u l t s from Figure (64) and (65) suggest t h a t t h e r e may be bubble nuc l ea t i on f o r mercury i n forced convect ion bulk bo i l i ng . I t i s a known f a c t t h a t a b o i l i n g mechanism based upon t h i s bubble nuc lea t i o n model p r e d i c t s t h e hea t t r a n s f e r coef f i c e n t t o be markedly in f luenced by hea t f l u x , v i r t u a l l y una f f ec t ed by ( 2 1) t h e vapor q u a l i t y , and increased wi th i n c r e a s i n g temperature leve 1 ( 2 2 ) Furthermore, t h e l i q u i d metal pool b o i l i n g d a t a of Bon i l l a , f o r which bubble nuc l ea t i on i s presumed t o occur , show an inc rease i n hea t t r a n s f e r c o e f f i c i e n t wi th both hea t f l u x and temperature l e v e l . On t h e con t r a ry , a b o i l i n g mechanism based upon t h e f i l m evapora t ion model p r e d i c t s t h e hea t t r a n s f e r c o e f f i c i e n t t o be increased wi th i n c r e a s i n g vapor q u a l i t y , decreased wi th i nc reas ing temperature level'''' , and, probably, in f luenced by heat f l u x , depending upon t h e p a r t i c u l a r a n a l y t i c a l flow model chosen. Therefore , t h e hea t f l u x and s a t u r a t i o n temperature ( o r s a t u r a t i o n pressure) a r e thought t o be important parameters a f f e c t i n g t h e heat t r a n s f e r c o e f f i c i e n t i n t h e nuc l ea t e b o i l i n g reg ion . The l o c a l vapor q u a l i t y i s presumed t o have no e f f e c t upon t h e hea t t r a n s f e r c o e f f i c i e n t i n t h i s r eg ion f o r a l l s a t u r a t i o n temperatures . ( 2 ) The mass ve loc i ty , 6, must a l s o be a s i g n i f i c a n t parameter a f f e c t i n g t h e hea t t r a n s f e r c o e f f i c i e n t . Phys i ca l l y , i n t h e nuc l ea t e b o i l i n g reg ion w i t h a t h i n l i q u i d l a y e r on t h e wa l l , t h e high flow r a t e would shea r o f f l i q u i d from t h i s annular l aye r t o i nc rease t h e hea t t r a n s f e r c o e f f i c i e n t . ( 3 ) The e f f e c t of t h e l o c a l a c c e l e r a t i o n f i e l d on t h e heat t r a n s f e r c o e f f i c i e n t i n t h e nuc lea te b o i l i n g reg ion is neglec ted . These t h r e e cons ide ra t ions r e s u l t e d i n t h e fol lowing func t iona l r e l a t i o n : a h = (cons tan t ) (G) NB ( 19) The expondent "c" was determined by t ak ing an average s lope from l i n e s 1, 2 & 3 presented i n F igure (58). The parameters G and PSat d id no t vary s i g n i f i c a n t l y among t h e s e l i n e s i n t h e de te rmina t ion of "c". Followiing t h i s approach, "c" was found t o be approximately 0.85. The exponent "b" h was then determined by p l o t t i n g "B ve r sus P on a log-log s c a l e a t ( q t t ) 0.85 S a t approximate cons tan t values of G. A s shown i n F igure (66) , "b" was thus determined t o be 0.54. S imi l a r ly , t h e exponent "a" was determined h from Figure (67) by p l o t t i n g NB versus G on a log-log (4joSs5 ('sat )0.54 s c a l e and was found t o be approximate1.y 0.65. F i n a l l y , from t h e c o r r e l a t i n g p lo t i n F igure (68) , t h e dimensional const an t was determined. The c o r r e l a t i o n equat ion has t h e fol lowing f i n a l form: - 4 h = 3.09 x 10 )O. 54(q,,)0. 85 Btu (Psa t - NB ( 2 0 ) hr- ftZ + with a s c a t t e r band of - 20%. Fur ther e f f o r t was made t o d e t e c t i f any t r end ex i s t ed i n h a s NB a func t ion of t h e l o c a l a c c e l e r a t i o n f i e l d generated by t h e b o i l e r i n s e r t . Figure (69) is a p lo t of t h e express ion h /(G) 0.65 NB ('sat )0.54 ( q t r ) O . 85 versus t h e l o c a l a c c e l e r a t i o n f i e l d , ( 1 + \) where A is R obtained from Equation (9 ) . The d a t a i n F igure (69) sugges ts t h a t t h e e f f e c t of t h e induced a c c e l e r a t i o n f i e l d upon h is e i t h e r NB n e g l i g i b l e o r is l e s s t han t h e d a t a s c a t t e r of about 2 20% over I h c range covered. C r i t i c a l Heat Flux, q"_ A s descr ibed i n Appendix ( I ) , t h e onset of t h e c r i t i c a l f l u x cond i t i on was determined by f i n d i n g an i n f l e c t i o n poin t a long t h e s h e l l temperature p r o f i l e . The c r i t i c a l hea t f l u x d a t a obtained i n t h i s manner are t abu la t ed i n Table ( 9 ) . Experiments i n pool b o i l i n g have ind i ca t ed t h a t t h e c r i t i c a l hea t f l u x is p ropor t i ona l t o t h e f o u r t h roo t of t h e l o c a l a c c e l e r a t i o n (25) (26) ; t h e r e f o r e , t h e r a d i a l a c c e l e r a t i o n produced by t h e h e l i c a l i n s e r t might l o g i c a l l y have a s i m i l a r e f f e c t on t h e c r i t i c a l hea t f l u x i n forced convect ion bo i l i ng . A s recommended i n Reference ( 6 ) , t h e c r t i c a l hea t f l u x d a t a can be c o r r e l a t e d i n t h e fo l lowing form, where q" is t h e c r i t i c a l hea t f l u x f o r pool b o i l i n g , and t h e empir ica l PBc const a n t , n , can be determined by experiment. References (5 ) and (6) c o r r e l a t e d forced convect i o n potassium c r i t i c a l heat f l u x d a t a reasonably w e l l by u s i n g t h e form of Equation (21) . The value of q" 6 2 was t aken t o be 10 Btu/ h r - f t , and PBc t h e cons t an t , n, was e m p i r i c a l l y determined t o be 2. Following t h i s procedure, t h e presen t Mercury c r i t i c a l hea t f l u x d a t a can be reasonably w e l l represen ted by t h e fol lowing equa t ion : 6 2 I n F igure (70) t h e c r i t i c a l hea t f l u x d a t a a r e presented by . . C: a p l o t of t h e express ion versus x t h e q u a l i t y a t c r i t i c a l c ' heat f l u x condi t ion . Also i n t f i i s f i g u r e , t h e b o i l i n g potassium c r i t i c a l heat f l u x d a t a from References (5) and (6) a r e p l o t t e d f o r compar i son. To determine t h e e f f e c t of b o i l i n g temperature and mass v e l o c i t y on t h e c r i t i c a l heat f l u x , Figure (71) was prepared. The r e s u l t shows t h a t t h e r e is no d e f i n i t e grouping of t h e d a t a wi th respec t t o e i t h e r mass v e l o c i t y o r b o i l i n g temperature. T r a n s i t i on Bo i l ing Region, h TB A s descr ibed before, a f t e r onset of t h e c r i t i c a l hea t f l u x cond i t i on and before establ ishment of t h e superheated vapor condi t ion , t h e t e s t s e c t i o n was i n t h e t r a n s i t i o n b o i l i n g region. This reg ion was d i s t i ngu i shed by t h e s h e l l s i d e temperature p r o f i l e a s shown i n F igures (21) t o ( 3 6 ) . I n t h e cons tan t heat f l u x b o i l i n g potassium e x ~ e r i m e n t s , ( ~ ) t h e t r a n s i t i o n b o i l i n g reg ion was observed by r e l a t i v e l y l a r g e tube w a l l temperature o s c i l l a t i o n s which increased i n amplitude when t h e q u a l i t y was increased. However, a s pointed out by Pe terson(6) i n h i s two-f l u id b o i l i n g potassium experiment, t h e s e temperature o s c i l l a t i o n s were l e s s i n magnitude i n t he t r a n s i t i o n b o i l i n g reg ion f o r t he two f l u i d case and t h e l o c a l h e a t f l u x decreased with inc reas ing q u a l i t y . Furthermore, as concluded i n Reference (6), the maximum p o s s i b l e w a l l temperature was l i m i t e d by the l o c a l primary f l u i d temperature, and i t was poss ib l e t h a t t he f i l m b o i l i n g reg ion was never reached. The t r a n s i t i o n b o i l i n g d a t a c a l c u l a t e d fol lowing t h e procedures i n Appendix ( I ) a r e presented i n Table (10) . An at tempt t o seek an empir ica l c o r r e l a t i o n of t h i s d a t a was made by cons ider ing t h e fol lowing a n a l y t i c a l flow model which is thought t o e x i s t i n t h e t r a n s i t i o n b o i l i n g region. The flow is assumed t o c o n s i s t of l i q u i d d r o p l e t s o r patches on t h e wa l l s u r f a c e posi t ioned i n a random manner wi th t h e remainder of t h e wa l l s u r f a c e t h a t is not covered by l i q u i d t o be exposed t o dry mercury vapor a t l o c a l s a t u r a t i o n temperatures . I f t h e heat t r a n s f e r of t h e en t r a ined mercury l i q u i d is neglec ted i n t h e vapor core , t h e heat t r a n s f e r mechanism f o r t h i s a n a l y t i c a l model can be assumed t o be l o c a l b o i l i n g f o r t h e l i q u i d drople t and forced convect ion f o r t h e dry vapor a r eas . Therefore, t h e average hea t t r a n s f e r r a t e can be w r i t t e n by combining t h e s e con t r ibu t ions as: 11 11 = q LB ave I I h = ave A~ ' + hvx TB - Tw T ~ a t *T where h i n d i c a t e s t h e vapor forced convect ion hea t t r a n s f e r c o e f f i c i e n t vx evaluated a t l o c a l condi t ions wi th f r a c t i o n a l mass of mercury vapor. M t e r n a t e l y , t h e equat ion can be r e w r i t t e n as : s i n c e A - + AV - A T . The s ingle-phase vapor forced convect ion hea t t r a n s f e r c o e f f i c i e n t evaluated a t l o c a l condi t ions can be simply ca l cu la t ed by t h e equation: wi th h given by t h e convent ional Di t tus -Boel te r o r Colburn equat ions v modified f o r h e l i c a l flow. The l o c a l l i q u i d b o i l i n g heat t r a n s f e r c o e f f i c i e n t , h can be approximately ca l cu la t ed by us ing t h e CichelPi- LB ' Boni l l a r e l a t i o n : where h is defined a s t h e heat t r a n s f e r c o e f f i c i e n t a t t h e c r i t i c a l LBC h e a t f l u x condit ion. For t h e purpose of s i m p l i f i c a t i o n , i f hVx Z h ~ " and h " LB - h ~ ~ ~ , then Equation (25) reduces t o L The problem remaining is t o eva lua t e t h e a r e a r a t i o - . I t is necessary Am 1 t o s p e c i f y t h e s i z e and d e n s i t y of t h e l i q u i d d r o p l e t s . To avoid A- L complicat ion i n t h i s r e s p e c t , two simple approaches t o eva lua t e - A~ a r e proposed. Case I Liquid Drople ts wi th Constant S i z e - I n t h i s ca se , t h e l i q u i d d r o p l e t s a r e a l l assumed t o have a disk-shape wi th diameter , d , and he igh t , d. The number of d r o p l e t s per u n i t a r ea , n, is assumed t o decrease when t h e d r o p l e t s s t a r t t o b o i l . Considering t h e l i q u i d volume w i t h i n a smal l tube length , AL, t h e fo l lowing equat ion i s obtained: where A = flow a r e a perpendicular t o tube a x i s . F Solving f o r n, Equation (29) yields: The r a t i o of l i q u i d a r e a t o t h e t o t a l hea t t r a n s f e r a r e a can thus be w r i t t e n as : Combining Equations (30) and (31) y i e lds : The assumed d rop le t diameter , d, can be e l imina ted from Equation (32) A, by us ing t h e cond i t i on a t & = d , and -2 = 1 (continuous l i q u i d f i l m C A~ up t o c r i t i c a l hea t f l u x po in t ) . Hence: Since the s i z e of t h e d r o p l e t i s assumed t o be unchanged, Equations (32) and (33) can be combined t o y i e l d a f i n a l express ion f o r t h e a r e a r a t i o : Case I1 Liquid Droplet with Constant Density - I n t h i s case , t h e number of l i q u i d d r o p l e t s per u n i t a r ea is assumed cons t an t , and t h e disk-shaped l i q u i d d r o p l e t s s h r i n k a s f l u i d evaporates . From equat ion (29) : Combining t h e above express ion i n t o Equation (31) y i e lds : A~ Using t h e condi t ion a t a = a and - = 1 gives: c ' Am F i n a l l y , s i n c e n is assumed t o remain unchanged a s t h e f l u i d evaporates , Equations (36) and (37) a r e combined y ie ld ing: A~ With t h e express ion i n Equations (32) and (36) f o r - , t h e f i n a l form A~ f o r c o r r e l a t i n g t h e t r a n s i t i o n b o i l i n g hea t t r a n s f e r c o e f f i c i e n t can be w r i t t e n a s : Case I Constant Droplet S i z e Case I1 Constant D r o ~ l e t Densitv '- can be w r i t t e n i n terms of l o c a l q u a l i t y and f l u i d The r a t i o -a, p r o p e r t i e s by us ing t h e c o n t i n u i t y equat ion f o r two-phase mixtures . This is: For mercury, t h e r a t i o Pv qL is n e g l i g i b l y smal l when compared wi th t h e va lues of one: I f a l s o , and i f K is assumed t o be cons t an t and n m e r i c a l l y equal t o K t h e equat ion can be s i m p l i f i e d to: c Consequently t h e p r e d i c t i o n s f o r h can be f u r t h e r s imp l i f i ed from TB Equations (39) and (40) . For t h e cons tan t d rop le t s i z e case: and f o r t h e cons tan t d r o p l e t d e n s i t y case: ~ h e s e two equat ions f i t two boundary po in t s of t h e t r a n s i t i o n b o i l i n g reg ion , i . e . , and h -h ( c r i t i c a l heat f l u x TB LB c condit ion) h .---;i~ h (superheated vapor TB s t a r t s ) Figure (72) shows a p l o t of versus ( l - X ) x c us ing LBc Vx ( 1 - xc)x t h e t r a n s i t i o n b o i l i n g d a t a l i s t e d i n able (10) . The values of h LBc were taken from t h e c r i t i c a l heat f l u x r e s u l t s - i n Table ( 9 ) , and h vx was ca l cu la t ed f o r mercury vapor a t l o c a l temperatures by t h e modified Di t tus -Boel te r Equation ( 2 ) . This f i g u r e shows a remarkably good c o r r e l a t i o n of t h e t r a n s i t i o n b o i l i n g da t a . Except f o r t h e regions near t h e two boundary po in t s , i . e . , c r i t i c a l hea t f l u x condi t ion and superheated vapor s t a r t i n g po in t , a l l t h e d a t a can be represented l i n e a r l y by t h e fol lowing empir ica l expression: TO s e e how good t h e a n a l y t i c a l p red ic t ions would be f o r h i n TB Equations (43) and (44) , Figures (73) and (74) were p l o t t e d us ing those equat ions toge the r wi th t h e present da t a . These f i g u r e s show t h a t t h e a n a l y t i c a l p red ic t ions y i e l d much higher values when compared t o t h e experimental r e s u l t s . Since t h e major assumption used i n t h e d e r i v a t i o n of t h e s e a n a l y t i c a l p red ic t ions was t h e n e g l i g i b l e e f f e c t from en t r a ined l i q u i d d r o p l e t s , t h e d i sc repanc ie s between t h e a n a l y t i c a l p red ic t ions and t h e experimental r e s u l t s might be a t t r i b u t e d mainly t o t h e l i q u i d entrainment e f f e c t . Following t h i s argument, a parameter, 0, was considered a s t h e entrainment e f f e c t and was def ined as : Entrainment a t any l o c a t i o n i n t h e t r a n s i t i o n b o i l i n g reg ion 0 = Entrainment a t c r i t i c a l f l u x cond i t i on Hence a s e r i e s of parametr ic l i n e s were added t o F igures (75) and (76) by in t roducing t h e parameter, 0, t o Equations (43) and (44) . The parameter S, shows t h e inadequacy of t h e a n a l y t i c a l flow models t o account f o r t h e a c t u a l p i c t u r e of t r a n s i t i o n bo i l i ng . The parametr ic l i n e wi th 0 = 1 .0 impl ies t h a t t h e l i q u i d entrainment throughout t h e t r a n s i t i on b o i l i n g reg ion is kept v i r t u a l l y unchanged, and i ts value is i n h e r i t e d from t h e c r i t i c a l hea t f l u x cond i t i on throughout t h e t r a n s i t i o n b o i l i n g region. I n t h i s way, a l l t h e d a t a p o i n t s were i d e n t i f i e d from t h e s e f i g u r e s a s having t h e i r mwn ind iv idua l fl values. A p lo t of 0 versus ( 1 + A ) f o r each d a t a poinl was then cons t ruc t ed , R a s shown i n F igure ( 7 5 ) , f o r both a n a l y t i c a l flow models considered. I t can be seen from Figure (75) t h a t t h e entrainment parameter,^, i s s t r o n g l y inf luenced by l o c a l r a d i a l a c c e l e r a t i o n s developed by t h e h e l i c a l i n s e r t . Moreover, t h e value of 0 seems t o approach i t s lower l i m i t a symptot ica l ly a s t h e value of ( 1 + AR) goes h igher . This r e s u l t tends t o confirm t h e i n t e r p r e t a t i o n of a s an entrainment e f f e c t , s i n c e t h e r a d i a l a c c e l e r a t i o n should reduce entrainment . The r e s u l t a l s o sugges ts t h a t vor tex i n s e r t s tend t o e l imina te t h e l i q u i d entrainment i n t h e t r a n s i t i o n b o i l i n g region. However, t h e present r e s u l t s only provide a q u a l i t a t i v e conclusion about t h e entrainment e f f e c t , and it i s be l ieved t h a t more experiments w i th var ious s i z e s of i n s e r t s should be c a r r i e d out i n t h i s a r e a before drawing a complete p i c t u r e of t h e entrainment behavior i n t h e t r a n s i t i o n b o i l i n g region. Empir ical express ions f o r 0 i n t e r m s of r a d i a l a c c e l e r a t i o n s .. were obtained from Figure (75) . For t h e cons tan t d e n s i t y model: and f o r t h e cons tan t s i z e model: I t should be not iced t h a t t h e expressions f o r 0 i n Equations (47) and (48) were obtained only on an experimental b a s i s and under t h e present t e s t condi t ions . Cor re l a t ions f o r Liquid metal i n t h e t r a n s i t ion boi l i n g reg ion were a v a i l a b l e from cu r ren t l i t e r a t u r e (5) ( 6 ) . A s suggested i n Reference ( 6 ) , t h e fol lowing express ion was used t o c o r r e l a t e potassium t r a n s i t i o n b o i l i n g da t a : (24) Following a conclusion from pool b o i l i n g r e s u l t s i n t h e g r a v i t y f i e l d , t h e exponent a was given values of 1/4 t o 1/5. B, b, and c were determined purely by the bes t f i t wi th t h e d a t a . A s shown i n F igure (76) a l l t h e d a t a grouped n i c e l y when represented by t h e fol lowing equat ion: I t must be emphasized, however, t h a t Equation (50) is p r imar i ly an empir ica l express ion covering only t h e present d a t a range. Any e x t r a p o l a t i o n ou t s ide t h e d a t a range should be made wi th caut ion . Superheated Vapor Region, h v The superheated reg ion i s def ined i n t h e mercury temperature p r o f i l e a s t h e r eg ion i n i t i a t e d by a sudden temperature r i s e . The l o c a l heat t r a n s f e r c o e f f i c i e n t s i n t h i s reg ion were obtained by the c a l c u l a t i o n a l procedure d iscussed i n Appendix ( I ) , and t h e l o c a l d a t a a r e summarized i n Table (11) . The convent ional s ingle-phase p red ic t ion , t h e Di t tus -Boel te r equat ion, was used f o r comparison wi th t h e da t a . To account f o r t h e e f f e c t of t h e vor tex i n s e r t , h e l i c a l parameters were introduced. The modified Di t tus -Boel te r equat ion(6) f o r h e l i c a l flow through tubes i s t h e r e f o r e : where t h e subscr ip t b ind ica tes evaluat ion a t l o c a l bulk temperature. A s discussed i n Appendix ( I ) , t h e h e l i x parameter f o r heat t r a n s f e r was obtained f o r Equation (51) by employing t h e following re la t ionship: Hence, Equation ( 51) was reduced t o : with h given, f o r a x i a l flow only, as: Va To account f o r t h e h e l i c a l flow e f f e c t , a heat t r a n s f e r parameter, , was obtained from Equation (53), a s follows: This is s i m i l a r t o the pressure parameter, K which was defined i n P ' Equation (A52) . Figure (77) shows a p lo t of t h e superheated vapor da ta together with t h e predic t ion from Equation (55). The r e s u l t s shows t h a t the modified Dit tus-Boelter equation with t h e heat t r a n s f e r parameter K is q u i t e adequate t o predic t the heat t r a n s f e r coe f f i c i en t h f o r mercury vapor flowing i n a tube with a h e l i c a l i n s e r t . However, t h e parameter K i n Equation (55) is good only f o r c e r t a i n types of h 73 i n s e r t s ; f o r i n s t ance , t h e use of Kh t o p r e d i c t h f a i l e d i n t h e VH wire c o i l i n s e r t r eg ion i n t h e presen t t e s t . The mercury temperature p r o f i l e s and t h e hea t t r a n s f e r d a t a f o r t h e wi re c o i l i n s e r t r eg ion (beyond t h e t ube l eng th of 12 f e e t ) a r e recorded i n Table (11) under Run Nos. 9, 10, 11, 12, 13 and 14. I t was found t h a t i n t roduc ing K a s expressed i n Equation (55) g ives h an underes t imate of t h e test r e s u l t s . S ince t h e r e is no a n a l y t i c a l express ion f o r K app l i cab l e t o t h e wi re c o i l i n s e r t , t h e only h eva lua t ion f o r K i s experimental . From Figure (78) which shows h r e s u l t s f o r heated a i r hea t t r a n s f e r c o e f f i c i e n t s w i th var ious types of i n s e r t s i n tubes(27) , an experimentally-determined hea t t r a n s f e r parameter was obtained f o r t h e p r e s e n t l y employed wire c o i l i n s e r t geometry (d/Di = 0.1428, P/D. = 0.97). The K va lue was found t o be 1 h approximately 2.6. The hea t t r a n s f e r d a t a i n t h e wi re c o i l r eg ion a r e shown i n Figure (79) t o g e t h e r w i t h Equation (53) (u s ing \ equal t o 2 .6) , and t h e agreement is q u i t e good. To show t h e d e v i a t i o n of flow w i t h a wi re c o i l i n s e r t from a x i a l flow, Equat ion (54) was a l s o p l o t t e d i n F igure (79) . Th i s r e s u l t i n d i c a t e s t h a t s i n g l e phase t e s t s w i th a i r can be employed t o p r e d i c t mercury superheated vapor heat t r a n s f e r c o e f f i c i e n t s , even without an e s t a b l i s h e d t h e o r e t i c a l forumulat ion. 3, Local Two-Phase P re s su re Drop The l o c a l s a t u r a t i o n pressures of b o i l i n g mercury a t t h e measured mercury s a t u r a t i o n temperatures were obtained from t h e vapor p re s su re curve, and t h e r e s u l t s were p l o t t e d i n F igures (21) t o (36) f o r a l l t h e test runs. The l o c a l two-phase p re s su re g r a d i e n t s i n t h e b o i l i n g reg ion , ob ta ined through Equation (A75), a r e t abu la t ed i n 74 Table (12) f o r var ious pos i t i ons along t h e b o i l i n g length . Figure (80) shows t h a t t h e l o c a l two-phase pressure g rad ien t v a r i e s along t h e normalized b o i l i n g length , i n t h e low and in te rmedia te q u a l i t y regions and becoming cons tan t i n t h e h igh q u a l i t y reg ion . & L In F igures (81) and (82) , t h e e f f e c t s of saturation temperature and mass ve loc i ty on t h e l o c a l v a r i a t i o n of a r e presented. The two- TP phase momentum p res su re g rad ien t s c a l c u l a t e d by us ing Equation (A77) a r e t a b u l a t e d i n Table (12) , and s e l c t e d values a r e p l o t t e d i n F igure ( f o r comparison. The two-phase f r i c t i o n a l , c a l c u l a t e d by s u b t r a c t i n g t h e moinentum pressure g rad ien t s TPf pressure g rad ien t from t h e t o t a l g rad ien t , a r e t a b u l a t e d i n Table (12) and represented i n Figures (83) t o (85). I n t h e s e f i g u r e s , t h e p red ic t ions from t h e modified Mart inel l i -Nelson model and t h e homogeneous model were a l s o p l o t t e d f o r comparison. The da t a f a l l gene ra l ly between' the two p red ic t ions i n t h e h igher q u a l i t y region, x ~ 0 . 3 . The genera l t r end of t h e present da t a , however, is c l o s e r t o t h e modified Mart inel l i -Nelson p r e d i c t i o n ; t h e homogeneous model g ives an underest imation of t h e two-phase f r i c t i o n a l pressure g rad ien t . The e l e v a t i o n pressure g rad ien t given i n Equation (A78) was found t o be a smal l f r a c t i o n of t h e t o t a l p ressure g rad ien t , p a r t i c u l a r l y i n t h e high q u a l i t y region. Hence, i ts con t r ibu t ion is neglected when computing from Equation (A76). Using t h e experiment a l l y-det ermined , t h e l o c a l two-phase p re s su re P f mu l t ip l e rc tde f ined i n Equation (A88) was c a l c u l a t e d , and t h e r e s u l t s a r e t abu la t ed i n Table (12) . ~ i ~ u r e s (86) t o (91) compare t h e s e experimentatlly-determined values w i th t h e modified Mart inel l i -Nelson model and t h e homogeneous model p red ic t ions , showing t h a t t h e present da t a of f a l l between t h e s e two p red ic t ions and show b e t t e r XP agreement wi th t h e modified Mart incl l i -Nelson model. In t h e low quality region, both models underestimate the actual test data. - The final conclusion drawn from the present results is that the pressure loss during forced convection boiling of mercury in trdbles containing helical inserts can be fairly well predicted by [the modi- fied Martinelli-Nelson model, provided the single-phase liquid pres- sure losses are known. Predictions for single-phase pressure drop through tubes containing helical inserts currently are available in References (28) and (29). C. Discussion Certain assumptions were made in reducing the data as described in the calculational procedures in Appendix (I). These assumptions are summarized here and their effects on the accuracy of the data are discussed. Assumptions (1) The outer shell wall temperature is taken as an average value over the measurements by the five calibrated thermocouples located at equal intervals circumferentially on the outer shell wall. (2) The axial NaK bulk temperature gradient is equal to the axial shell wall temperature gradient in the boiling region. (3) The Dwyerk equation (Equation A5) for calculating the shell- side NaK film coefficient is valid. (4) Effects of the axial conduction in the shell-side NaK flow and in both the tantalum tube and stainless steel NaK layer contaiment tube walls are negligible. (5) Effects of uneven thickness of the boiler tube wall, MaK layer, and NaK containment tube wall are neglected. (6) Axial variation of heat flux on the shell-side film co- efficient is neglected. (7) Heat transfer through the NaK layer is considered as conduction only. Hence, the effect from natural convection of this thin layer (0.0275-inch) is neglected, (8) Eccentricity of the shell and tube alignment and its effects on data measurement are not considered. Validity of Assumptions The validity of these assumptions was carefully examined. It was found that the major uncetainty arose from the prediction of the shell-side film coefficient by using Dwyer's equation. ~wyer's pre- diction for concentric annuli (Equation A5) is based upon the condi- tions (1) uniform heat flux at the inner wall, (2) fully developed temperature and velocity profiles, (3) negligible axial conduction, and (4) no effect of transverse temperature variation on the physical properties of the flowing metal. Of course, none of these conditions exist in the shell-side flow of the present two-fluid test. It is possible to calculate the axial variation in the local heat transfer coefficient for any arbitrary axial variation in wall heat flux, but the calculations are difficult and, consequently, few results are available. For the case of turbulent flow of low-Prandtl-number liquids, there are even fewer analytical predictions available to account for the variable wall heat flux effect. Hsu (30) carried out computations for the case of sinusoidal wall heat flux distlribution and slug-flow condition in pipes. It is believed that the slug-flow1 re- sults in Reference (30) can be used to estimate the axial variation of the heat transfer coefficient for turbulent flow if the Peclet number is not too large (< 500). As shown in Figures (41) to (561, 77 t h e p r e s e n t l iea t Flux d i s t r i b u t i o n f o r s e v e r a l t e s t r u n s were rosghly approximated hy s i n u s o i d a l d i s t r i b u t i o n , exc lud ing t h e end r e g i o 7 s . F i g u r e ( 9 2 ) , a s c i t e d from IIsuls r e s u l t s , shows t h e a x i a l va , r ia t l lon of t h e l o c a l K u s s e l t number a s a f u n c t i o n of t h e n o n - d i ~ ~ e n s i o n a l p i ~ e number v a i u e is t h e same a s t h a t f o r uniform-wall h e a t f l u x , b e c a u s e t h e h e a t f l u x i n t h a t r e g i o n i s r e l a t i v e l y uniform. Excluding t:e end r e g i o n s , t h e a x i a l v a r i a t i o n of t h e l o c a l N u s s e l t nunber d i r .Lnishes 'peD s D a s t h e p a r a m e t e r - L d e c r e a s e s . For example, cu rve So. 1 (9 = 2.0) A d shows t h a t t h e l o c a l N u s s e l t number cou ld b e r e p l a c e d by t h e v a l u e c a l c u l a t e d f o r uniform-wall h e a t f l u x ( a t z/L = 0.5) o v e r most o f t h e p i p e l e n g t h w i t h o u t producing any a p p r e c i a b l e e r r o r s . F o r t h e p r e s e n t N ~ e D t e s t where t h e pa ramete r - L is e q u a l t o 1 .52, t h e d e g r e e o f u n c e r t a i n t y i n c a l c u l a t i n g t h e s h e l l - s i d e N a K f i l m c o e f f i c i e n t t h r o u g h Rwyes's e q u a t i o n (based on uniform-wal l t e a t f l u x ) i s a p p r o x i s a t e i y + 87; c v e r - most of t h e a x i a l l e n g t h . A d d i t i o n a l a n a l y t i c a l e s t i m a t e s were made which v e r i f y t h i s con- c l u s i o n . X Nu p r o f i l e p l o t ana logous t o t h a t i n F i g u r e ( 9 2 ) i s n o t known t o Se a v a i l a b l e f o r any p r e s c r i b e d h e a t f l u x d i s t r i b u t i o n and t u r b u l e n t f low c o n d i t i o n s . ~ t e i n ' ~ ' ) h a s o u t l i n e d a p r o c e d u r e i c r d e t e r m i n i n g t h e l o c a l N u s s e l t n m b e r f o r f low th rough p i p e s , anr,i:l.i, and p a r a l l e l p l a t e s , where t h e v e l o c i t y p r o f i l e i s e s t a b l i s h e d aial where t h e w a l l h e a t f l u x may v a r y a s any a r b i t r a r y f u n c t i o n o f the a : x i a l d i s t a n c e . For r e l a t i v e l y long c h a n n e l and a x i a l d i s t a n c e n o t Eoo c l o s e t o t h e end r e g i o n s , S t e i n deve loped t h e e q u a t i o n : where $-W-H-F = fully developed heat transfer coefficient for uni- form wall heat flux. The coefficient R generally is a function of n E ~ r ' NRes and channel geometry. Unforunately, Stein's work is only verified for the infinite flat-plate case for which Rn can be computed from the relationship: The first five values of the coefficients Cn, and A' are given m in Reference (31) for turbulent flow. From Equation (56) we get: - (N~u) U-W-H-F (N~u) 00 local dn v 1 n=l d ( 2 1 ~ ~ ) ~ Equation (58) was used in several runs of the present test to check the effect of axial variation of heat flux upon shell-side NaK conductance. In general, the results were consistent with the trend indicated by curve No. 1 in Figure (92). Clearly, the computed Equation (58) is not exactly for an annulus, in which case the co- 2 efficients C and A are not available at this time. It is felt, how- m m ever, that Equation (58) does indicate how the local Nusselt number would vary with axial length for the annular case since it is known that the infinite flat-plate case is one of the limiting cases of an annulus. Recent work on double h e a t exchangers a l s o provides u s e f u l in - formation on the p r e d i c t i o n of l o c a l h e a t t r a n s f e r c o e f f i c i e n t s . liunge and i n a p ioneer ing a n a l y t i c a l s tudy c a l c u l a t e d t h e h e a t t r a n s f e r c h a r a c t e r i s t i c s f o r laminar counter c u r r e n t flow i n double-pipe h e a t exchangers. They based t h e i r a n a l y s i s on a f u l l y developed flow condi t ion and assumed n e g l i g i b l e a x i a l conduction i n both t h e h e a t exchanger w a l l s and f l u i d s t reams. They obta ined l o c a l and asymptot ic Nusse l t number changes as func t ions of s e v e r a l s tandard parameters. The r e s u l t s i nd ica t ed t h a t f o r long exchangers, say L - L > 0.5, ( i n t he p re sen t t e s t - = N D 0.67) a fully-developed N D Pe e Pe e - U-w-H-F c o e f f i c i e n t may be used along almost t h e e n t i r e a x i a l length. However, Nunge and G i l l ' s r e s u l t s a r e only t r u e f o r laminar flow and a l l l i q u i d s . I n t h e case of l i q u i d meta l , f o r t u rbu len t flow a s f o r laminar flow, t h e l o c a l hea t t r a n s f e r c o e f f i c i e n t i s very much dependent on the thermal boundary condi t ions . For t h i s reason, t h e r e s u l t s drawn f o r laminar flow of any l i q u i d s i n Reference (32) can be appl ied t o t u rbu len t flow of l i q u i d metal. I n t h e absence of t u r b u l e n t flow r e s u l t s , i t i s probably appropr i a t e t o assume t h a t t h e turbulen t - flow c o e f f i c i e n t normalized wi th r e spec t t o t h e fully-developed tu rbu len t flow U-W-II-F c o e f f i c i e n t is t h e same a s t h e comparable normalized c o e f f i c i e n t f o r t h e laminar f law case. Summarizing a l l t h e foregoing d i scuss ions on t h e v a l i d i t y of Dwyer's equat ion t o c a l c u l a t e NaK f i l m c o e f f i c i e n t , i t is f e l t t h a t t h e degree of unce r t a in ty on h would approximately be 8%. This NaK unce r t a in ty is probably g r e a t e r i f t h e a x i a l conduction i n t h e s h e l l - s i d e s t ream and e c c e n t r i c i t y a r e p re sen t . 80 Estimates were also made of the axial tube wall conductions and were found to be negligible. No estimation was made on the heat transfer mechanism through the very confined static-NaK layer with an average thickness of 0.0275- inch. It is conceivable that some unstable, free-convection cells might set up and some temperature fluctuations might prevail at any axial and lateral positions. However, the complex flow patterns and thermal conditions which might exist in this layer generally preclude the feasibility of any analytical attack. The assumption of the axial shell wall temperature gradient in calculating local heat flux in the boiling region was justified through several heat balance calculations. The results generally checked within 5%. Finally, it is believed that the average shell wall temperature taken over the five wall thermocouples would definitely eliminate part of the effect coming from eccentricity of the tube-and-shell alignment. Accuracy of Data Local Heat Flux, q" - The degree of accuracy on the local heat flux calculation depends upon the assumption of the local shell wall temperature gradient and the way to obtain it. Through occasional heat balance checks and very careful performance on graphical dif- ferentiation in obtaining the local temperature gradient, it is be- lieved that the accuracy on local heat flux is approximately 6%. Local Overall Heat Transfer Coefficient, U - The accuracy of U, obtained from the local heat flux and the shell- and tube-side 8 1 fluid temperatures, depends primarily on the accuracy of those values. It is estimated that the present data for U may have a probable error of 10%. The local mercury saturation temperature required is the cal- culation for U is deterrhined from the temperature profile measured by insert thermocouples. The helical insert employed in the present test induces a swirling motion in the flow and imposes radial acceleration on the fluid. If the flow pattern is visualized as a relatively thin layer of liquid adjacent to the wall with a vapor core, then there exists a radial pressure difference between the fluid close to the wall and that at the centerline. If it is assumed that the vapor is saturated at the centerline then the centerline temperature (measured by insert thermocouples) must be corrected to obtain the saturation temperature close to the wall for use in the calculation of U. Such corrections can be made through using the following equation from Reference (5) : However, it is found that such corrections are generally not necessary when evaluating the local overall heat transfer coefficient U. Combined Wall Conductance, Ucw - The combined wall conductance, including those thermal resistances from the NaK film coefficient, stainless steel tube wall, static-NaK layer, and tantalum tube wall, has a variety of implicit errors. The major portion of error (approxi- mately 10%) comes from the method of predicting hNaK as previously discussed. No analytical estimates were made on tube and annuli wall thermal resistances, as they may be effected by their uneven wall thicknesses. Also, no estimate was made on the natural convection 8 2 e f f e c t on t h e h e a t t r a n s f e r r a t e through t h e very t h i n static-NaK l aye r . The r e l a t i v e thermal r e s i s t a n c e s f o r t h i s composite s t r u c t u r e a r e presented i n F igure (93) and t h e one f o r t h e static-NaK l a y e r w a s obtained by t r e a t i n g i t a s conduction only. A s t h e f i g u r e shows t h e K a K f i lm c o e f f i c i e n t r e s i s t a n c e occupied almost of t h e t o t a l combined r e s i s t a n c e s . The r e s i s t a n c e of t h e static-NaK l a y e r , i n which t h e n a t u r a l convect ion e f f e c t was no t considered, on ly occupied about 15% of t he t o t a l r e s i s t a n c e . I n t h i s r e spec t , unce r t a in ty i n p r e d i c t i n g the static-NaK thermal r e s i s t a n c e w i l l no t apprec iab ly a f f e c t t he accuracy of U Allowing a d d i t i o n a l unce r t a in ty due t o t hese unknown cw' e f f e c t s , t h e probable e r r o r i n U is es t imated t o be 15%. cw Local Mercury Heat T rans fe r C o e f f i c i e n t , h - The mercury h e a t Elg t r a n s f e r c o e f f i c i e n t s f o r va r ious reg ions throughout t h e b o i l e r were obtained by s u b t r a c t i n g t h e combined thermal r e s i s t a n c e U from t h e cw o v e r a l l r e s i s t a n c e U c a l c u l a t e d from l o c a l h e a t f l u x and temperature measurements. Iience , t h e accuracy of hHg w i l l va ry f o r d i f f e r e n t hea t t r a n s f e r reg ions . I n F igure (94) t y p i c a l d a t a of h%' are sented and t h e s u b t r a c t i o n performed t o o b t a i n t h e s e l o c a l v a l u e s of h i s a l s o i l l u s t r a t e d . Consequently, a d e t a i l e d e r r o r a n a l y s i s Hg can be performed on h by us ing Figure ( 9 4 ) and t h e probable m m i m m Hg e r r o r s f o r U and U discussed before. I n gene ra l , t h e a n a l y s i s in- CW d i ca t ed t h a t no s e r i o u s e r r o r ( l e s s than 15%) can be made i n h ob- hz t a ined i n t h e t r a n s i t i o n , subcooled l i q u i d , and superheated vapor reg ions , s i n c e over t h e s e reg ions t h e mercury thermal r e s i s t a n c e is a r e l a t i v e l y l a r g e f r a c t i o n of t h e o v e r a l l h e a t t r a n s f e r c o e f f i c i e n t . However, f o r t h e n u c l e a t e b o i l i n g reg ion , a s shown i n F igure ( 9 4 ) , t h e accuracy on h dec reases a s h is taken from l o c a t i o n s c l o s e r IIg IIg t o t h e c r i t i c a l hea t f l u x po in t where t h e thermal r e s i s t a n c e i s small . 83 The worst c a s e might happen a s t h e expected e r r o r s i n t h e p r e d i c t i o n of U a r e l a r g e r than t h e ind ica t ed mercury hea t t r a n s f e r c o e f f i c i e n t C W taken c l o s e t o t h e end of t h e n u c l e a t e b o i l i n g region. Typical ex-, 2 amples showed t h a t when h was 5500 Btu/hr-f t -OF o r l a r g e r , t h e Hg est imated e r r o r f o r h might exceed 35% i f a 10% e r r o r f o r U and a Hg 15% e r r o r f o r U were assumed. For t h i s reason, when e s t a b l i s h i n g C W t h e c o r r e l a t i o n equat ion f o r h i n t h e n u c l e a t e b o i l i n g reg ion , Hg h va lues over 5500 ~ t u / h r - f t ~ - ~ ~ were not employed. Ilg C r i t i c a l Heat Flux, q" - The c r i t i c a l hea t f l u x d a t a is be l ieved C more accu ra t e than h d a t a and is est imated t o have a probable e r r o r Hg of 15%. The accuracy of q" pr imar i ly depends on how we l l t h e l o c a l C g r aph ica l d i f f e r e n t i a t i o n is performed a long t h e s h e l l temperature p r o f i l e and on how a c c u r a t e t h e c r i t i c a l hea t f l u x p o i n t is loca t ed . V CONCLUDING REMARKS A cons ide rab l e amount of experimental and a n a l y t i c a l in format ion on forced convect ion b o i l i n g of mercury i n a tan ta lum tube has been obtained from t h e p re sen t i n v e s t i g a t i o n , The r e s u l t s p resen ted i n t h i s r epo r t inc lude both l o c a l and average hea t t r a n s f e r and p re s su re drop d a t a from subcooled l i q u i d t o superheated vapor cond i t i ons f o r mercury a t s a t u r a t i o n temperatures from 975OF t o 1120°F. The experimental d a t a and t h e a s soc i a t ed c o r r e l a t i o n s a r e d i r e c t l y a p p l i c a b l e t o t h e des ign of once-t hrough b o i l e r s f o r Space Power Rankine Cycle Systems employing mercury a s a working f l u i d . S p e c i f i c conc lus ions wi th r e spec t t o t h e heat t r a n s f e r and p re s su re drop of b o i l i n g mercury follow. A. Heat T rans fe r (1) Liquid mercury hea t t r a n s f e r c o e f f i c i e n t s , h ~ 7 were found t o 2 b e i n t h e range of 980 t o 1850 Btu/hr-f t -OF. They a r e i n f a i r agreement- with t h e Lubars ky-Kaufman equat ion ' 15)when t h e modified Pec l a t G D C Number, = , is less t h a n 250, and wi th Lyon's equa t ion ( 14) N ~ e k .when P is g r e a t e r than 250. Both equa t ions were eva lua ted u s ing h e l i c a l e flow parameters. (2) The mercury r iucleate b o i l i n g hea t t r a n s f e r c o e f f i c i e n t , %B: 2 was found t o be i n t h e range of 2100 t o 7050 Btu/hr-f t -OF. General ly , t hey i n c r e a s e wi th increased hea t f l u x (F igure 60) and a r e v i r t u a l l y independent of q u a l i t y (F igure 61) . However, f o r q u a l i t y l e s s t han 0.15, h i nc reases s l i g h t l y wi th i nc reas ing q u a l i t y . NB ( 3 ) Vortex genera tor i n s e r t s were found t o i nc rease h f o r t h e m q u a l i t y range l e s s t han 0.10. I n t h e q u a l i t y range 0.10 < x < x t h e c ' i nc rease i n h duc t o i n s e r t i s g r a d u a l l l y diminished. NB (4) The heat t r a n s f e r mechanism i n t h e nuc l ea t e b o i l i n g reg ion i s be l ieved t o be bubble nuc lea t ion . Th i s conc lus ion is supported by t h e f a c t t h a t t h e measured h i n c r e a s e s wi th i n c r e a s i n g hea t f l u x and i s NB v i r t u a l l y independent of q u a l i t y (F igures 60 and 61). A q u a l i t a t i v e s tudy concerning t h e hea t t r a n s f e r model a l s o a f f i rmed t h i s conc lus ion (F igures 64 and 65). ( 5 ) Empir ical c o r r e l a t i o n of h (F igure 68) shows t h a t 11 NB NB i nc reases a s t h e mass v e l o c i t y and s a t u r a t i o n p re s su re i nc rease and approximately i n propor t ion t o t h e 0.85 power of t h e heat f l u x under t h e presen t t e s t cond i t i ons . (6 ) The c r i t i c a l hea t f l u x q" was found t o be i n t h e range of C 5 5 2 1.86 x 10 t o 4 .5 x 10 Btu/hr-f t . The va lue of q" was found t o C decrease wi th i n c r e a s i n g l o c a l vapor q u a l i t y (F igure 70) and t o i nc rease approximately i n p ropor t i on t o t h e f o u r t h roo t of t h e l o c a l a c c e l e r a t i o n f i e l d . (7) E f f e c t s of mass v e l o c i t y and s a t u r a t i o n temperature on t h e c r i t i c a l hea t f l u x were not observed w i t h i n t h e presen t range of experimental d a t a . 8 6 (8) The transition boiling heat transfer coefficient, hTB, was found to be in the range of 1530 to 6500 ~tu/hr-ftL-OF. The values of h~ B generally decrease with increasing quality or with decreasing heat flux. ( 9 ) Local data of hTB were successfully correlated by the equation suggested in Reference (6). The correlation (Equation 50) states t h a t hTB decreases with increasing vapor quality, increases with increasing local acceleration field, and decreases with increasing wall to fluid temperature difference. (10) Liquid droplet entrainment was suggested to be significant in the transition boiling region and to be strongly dependent upon the local acceleration field developed by the insert (Figure 75). (11) An analysis based on adopting a flow model of constant droplet number in the transition boiling region was successful in correlating the present transition boiling heat transfer data after introducing the entrainment parameter (Figure 72). (12) The large differences between wall and fluid temperature which characterize the film boiling region were not observed in the prescmt two-fluid, temperature-controlled experiment. (13) The superheated vapor heat transfer coefficient, hV, is in 2 the range of 45 to 110 Btu/hr-ft -OF. The values can be generally correlated in the helical insert region by the conventional Dittus-Boelter equation after introducing the helical flow heat transfer parmeter, Klh (Equation 55). (14) I n t h e case of t h e wi re c o i l i n s e r t , t h e Di t tus -Boel te r equat ion f a i l s t o c o r r e l a t e t h e test d a t a f o r hV even a f t e r being modified f o r h e l i c a l flow. The h e a t t r a n s f e r parameter, Kh, w a s found t o be 2.6 f o r the present wi re c o i l i n s e r t geometry, P/Di = 0.97 and d/Di = 0.1426, agreeing wi th measurements made wi th a i r repor ted i n Reference (27). B. P ressure Drop (1) Single-phase p re s su re drop f o r t h e h e l i c a l i n s e r t can be predic ted wi th reasonable accuracy by us ing t h e convent ional Fanning equat ion a f t e r in t roducing a p re s su re parameter, K f o r t h e h e l i c a l P' flow condi t ion . (2) Single-phase p re s su re drop f o r t h e wi re c o i l i n s e r t can no t be pred ic ted by t h e same method a s t h e h e l i c a l i n s e r t . The p re s su re parameter, K was found t o be 23 + 22% f o r t h e p re sen t w i r e c o i l i n s e r t P ' - geometry P/D = 0.97 and d/Di = 0.1428, agree ing roughly wi th measure- i ments made wi th a i r repor ted i n Reference (27). ( 3 ) Local two-phase p re s su re g r a d i e n t s were found t o be dependent upon the b o i l i n g length . They i n c r e a s e r a p i d l y as t h e b o i l i n g starts and g radua l ly become cons tan t as t h e end of t h e b o i l i n g reg ion i s reached. (4) Local two-phase momentum p res su re g r a d i e n t s were found t o be only about 10 percent of t h e t o t a l two-phase p re s su re g rad ien t s . tIeace, t he f r i c t i o n a l p re s su re g r a d i e n t s dominate t h e p re s su re drop i n t h e two- phase region. 8 8 (5) Local two-phase pressure gradients were found to fall generally between the prediction of the modified Martinelli-Nelson model and the homogeneous model. (6) In the low quality region, both predictions give an underestimate of the two-phase frictional pressure gradient data. ( 7 ) The two-phase frictional pressure gradients were found to increase with decreasing saturation temperature in the low quality region and to decrease with decreasing saturation temperature in the high quality region. (8) Experimental results show that the two-phase frictional pres- sure drop multiplier can be predicted with reasonable accuracy by using the Martinelli-Nelson correlation modified for mercury. The homogeneous model, though simpler in use, was found to be inadequate to predict the multiplier. (9) Effects of mass velocity on the two-phase pressure drop multi- plier were apparently not significant in the present test ranges. A S h e l l T i C (Average Value) a,$) Well T,;C () insert T ~ C Saturat ion Temp. v i a . Taylor Gauge c a l c u l a t i o n v i a E q . (A701 v Satural:ion Pressure v i a . Inser t T/'C Taylor Gauge DATE 12-29-67 TIME 1350 w 0.1995 lb/sec S t a t i o n # A s h e l l Tic (Average Value) @ I n s e r t T ~ C d, C a l c u l a t i o n v i a Eq. (A70) Tay lor Gauge Wcll T.;C (O Sat.uration Temp. v i a . Taylor Gauge Saturat ion Pressure v ia . I n s e r t T/'C 0 2 4 6 8 10 12 14 16 Distance from N a K Exit, feet RUN N o . 4 DATE 12-29-67 TIME 2005 NaK S i d e Hg S i d e Qnet 54.43 Pin 334.2 "IA Pex 261.8 PSIA AP 72.4 PSI Tes 1060 OF ATsh 206.6 OF T i n 760.4 OF Tex1310 OF Ti s 1107 OF F i g u r e 24. SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . s h e l l T/C (Average Value) @ I n s e r t T/C a C a l c u l a t i o n v i a Eq . (A70) T a y l o r Gauge Q,O Well T C C) S a t u r a t i o n Temp. v i a . T a y l o r Gauge \7 Safut'al ion P r e s s u r e v i a . I n s e r t T C DATE 1 2 -29-67 T I 2255 NaK S i d e W 2 . 0 4 6 lb/sec W 0 . 4 3 5 Ib/S@c F i g u r e 27 . SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . Shel l T/C (Average Value) 0,O Wcll T,'C @ Insert T'C Calculation v ia Eq. (~70) Taylor Gauge @ Saturation Temp. v i a . Taylor Gauge v Sat nratlon Pressure v i a . Insert Tic 8€ia aoo 1 0 2 4 6 8 10 12 14 16 Distance from NaK E x i t , feet RUNNo. 8 DATE 12-29-67 TIME 2315 w 2.013 lb/sec W 0.3987 lb/sec Figure 28. SNAP-8 S ing l e Tube B o i l e r Temperature and P re s su re P r o f i l e . S h e l l T/C (Average Value) @ Insert T ,C A Calculat ion v i a Eq . (A70) Taylor Gauge 0,0 Well T:'C C) Saturation Temp. v i a . Taylor Gauge v Sat urai: ion Pressure v i a . Insert T,.'C RUN No. 9 NaK S i d e W 1 . 9 1 8 lb/sec W 0 . 4 6 8 l b / s e c Dp 55 .8 PSI 1)L S h e l l T/C (Average Value) @ Insert T ~ C c a l c u l a t i o n v i a Eq. (A70) T a y l o r Gauge 0.0 Well T C @ Saturation Temp. v i a . T a y l o r Gauge r;r Salrrration P r e s s u r e v i a . Insert T;C 700 0 2 4 6 8 10 12 14 16 Distance from NaK Exit , feet DATE 1-3-68 TIME 1300 W 0 . 4 5 2 l b / s e c 250 .8 PSIA F i g u r e 30. SNAP-8 S i n g l e Tube B o i l e r Temperature and P re s su re P r o f i l e . 1) Shell TIC (Average Value) Q) Insert T'C A Calculation v i a ~ q . (A70) 7 Taylor Gauge 0,0 Well T.'C @ Saturation Temp. via . Taylor Gauge v Ssf ural.ion Pressure v ia . Insert Ti'C RUN N o . 12 W 1 .885 Lb/sec W 0.4718 lb /sec 345.6 PSIA 2 4 5 7 8 10 11 13 14 16 17 19 20 21 22 23 Shell S t a t i o n # 700 0 2 4 6 8 10 12 14 16 D i s t a n c e from NaK Exit, f e e t F i g u r e 32. SNAP-8 S i n g l e Tube B o i l e r Tempera tu re and P r e s s u r e P r o f i l e . A S h e l l T/(: ( A V ~ ~ R ~ C Value) @ I n s e r t T I C a C a l c u l a t i o n v i a Eq. (A70) Tay lor Gauge 0,0 Well T;C Q Suturat ion Temp. v i a . Tay lor Gauge v Salurat ion Pressure v i a . I n s e r t T C 2 4 5 7 8 10 11 13 14 16 17 19 20 21 22 23 Shell St 700 D i s t a n c e from NaK E x i t , feet RUN No. 13 DATE 1-6-68 2300 NaK S i d e H g S i d e W 0.47 l b / s e c Qnet 68.32 KW Pin 357.2 PSIA Pex 267.8 PS IA ap 89.4 PSI Tes 1067 OF ATsh 190 OF Tin 660 OF Te, 1291 OF Ti s 1103 OF a t i o n X F i g u r e 33. SNAP-8 S i n g l e Tube B o i l e r Temperature and P r e s s u r e P r o f i l e . 1)) S h e l l T/C (Average Value) @ I n s e r t T 'C C a l c u l a t i o n VIR Eq. (A70) T a y l o r Gauge 0,0 Wcll T C @ S a t u r a t i o n Temp. v i a . T a y l o r Gauge S e i u r a i i o n P r e s s u r e v i a . I n s e r t T C - DATE 1-9-68 TIME 5 1 5 NRK S i d e W 0 .731 l b / s e c Figure 35. SNAP-8 Single Tube Boiler Temperature and Pressure Profile. A S h e l l T/C (Average Value) @ Inser t T'C A Calcul .at ion v i a Eq. (A70) Taylor Gauge 0,0 Wel.1 T;'C @ Sat ,urat ion Temp. v i a . Taylor Gauge Sat:urat:ion Pressure v i a . I n s e r t T.;C DATE 1-9-68 TIME 2200 NaK S i d e W 0.351 lb/sec 2 4 5 7 8 10 11 13 14 16 17 19 20 21 22 23 Shell S t a t i o n # 700 0 2 4 6 8 10 12 14 16 D i s t a n c e from N a K E x i t , f e e t F i g u r e 3 6 . SNAP-8 S i n g l e Tube B o i l e r Tempera tu re and P r e s s u r e P r o f i l e , 6 + a rn El k a) 0 3 0.4 n cd Cd 0 a > > 4 . e w ar a m a r Distance from b o i l e r i n l e t , L, f t . Figure 41. Local Values of Heat Flux and Overa l l Heat Transfer Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 110 Run - 2 C C r i t i c a l Heat Flux Point IS I n i t i a t i o n of Superheat d) Calculated By Equation ( 2 ) -3 2 . 1 - _ - - - - i \ ! Distance from b o i l e r i n l e t , L, f t . Figure 42. Local Values of Heat Flux and Overa l l Heat Transfer Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 111 Run - 3 .IB Inception of Boil ing C C r i t i c a l Heat Flux Point IS I n i t i a t i o n of Superheat @ Calculated By Equation ( 2 ) C - - - - - - - - \ \ \ \ ', \ \ \ lo2 0 2 4 6 8 10 12 14 16 I8 Distance from b o i l e r i n l e t , L, f t . Figure 4 3 . Local Values of Heat Flux and Overall Heat Transfer Coefficient f o r Mercury Once-Through Boiling Runs. 112 Distance from boiler inlet, L , f t . F igure 44. Local Values of Heat Flux and Ove ra l l Heat Transfer Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. Distance from boiler inlet, L, ft. Figure 45. Local Values of Heat Flux and Overa l l Heat T rans fe r Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 114 IB Inception of Boi l ing C C r i t i c a l Heat Flux Point I S Initiation of Superheat @ Calculated By Equatlon ( 2 ) 7 d " 1 2 0 2 4 6 8 10 12 14 16 18 Distance from b o i l e r i n l e t , L, f t . F igure 46. Local Values of Heat Flux and Ove ra l l Heat T rans f e r Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 115 lo6 , Run - 7 IB Inception of Boil ing C Crit ical Heat Flux Point IS In i t ia t ion of Superheat @ Calculated By Equation ( 2 ) Distance from boi ler i n l e t , L, ft. Figure 47. Local Values of Heat Flux and Overall Heat Transfer Coefficient for Mercury Once-Through Boiling Runs. 116 Distance from boiler inlet, L , f t . F igure 48. Local Values of Heat Flux and Ove ra l l Heat T rans f e r Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 117 I B Inception of Boil ing C C r i t i c a l Heat Flux Point IS I n i t i a t i o n of Superheat @ Calculated By Equation ( 2 ) 7 - I Distance from b o i l e r i n l e t , L, f t . F igure 50. Local Values of Heat Flux and Overa l l Heat Transfer Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 119 I B Inception of Boi l ing C C r i t i c a l Heat Flux Point IS I n i t i a t i o n of Superheat (fg) Calcula t ed By Equation ( 2 ) 0 2 4 6 8 10 12 14 16 18 Distance from b o i l e r i n l e t , L , f t . Figure 51. Local Values of Heat Flux and Overa l l Heat T rans fe r Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 120 Distance from b o i l e r i n l e t , L, f t . Figure 52. Local Values of Heat F lux and Overall Heat Transfer Coefficient f o r Mercury Once-Through Boiling Runs. 121 Distance from b o i l e r i n l e t , ~ , f t . Figure 53. Local Values of Heat Flux and Overal l Heat Transfer Coeff ic ien t f o r Mercury Once-Through Boil ing Runs. 122 Distance from boiler inlet, L , f t . Figure 54. Local Values of Heat Flux and Overall Heat Transfer Coefficient for Mercury Once-Through Boiling Runs. 123 RUN - 15 IB Incept ion of Bo i l i ng C C r i t i c a l Heat Flux Point IS I n i t i a t i o n of Superheat @ Calcula ted By Equation ( 2 ) Distance from b o i l e r i n l e t , L, f t . Figure 55. Local Values of Heat Flux and Overa l l Heat T rans fe r Coe f f i c i en t f o r Mercury Once-Through Boi l ing Runs. 124 RUN -16 IB Inception of Boil ing C C r i t i c a l Heat Flux Point IS I n i t i a t i o n of Superheat @ Calculated By Equation ( 2 ) ----- Distance from b o i l e r i n l e t , L , f t . Figure 56. Local Values of Heat Flux and Overa l l Heat Transfer Coeff ic ien t f o r Mercury Once-Through Boil ing Runs. 125 h Fr 0 ~ 0 0 0 0 o o a v) U N N O m a h a - ( . - I - l - ( o o o m 4 d E-c Pool B o i l i n g , Ref. ( 1 7 ) Poo l B o i l i n g a t C r i t i c a l Heat Flwc L e v e l s Forced Convect ion, Eq ( ~ b l s e c - ~ t ~ ) (OF) 0 192 1120 A 185 1120 1 7 7 1110 @ 158 1100 1 4 1 1090 Ref. (16) 1 0 20 30 40 50 100 0 Temperature D i f f e r e n c e , AT = Tw - Tsats F F i g u r e 59. Mercury Nuc lea te B o i l i n g Hea t T r a n s f e r Data From a 0.67-Inch I D Tube With H e l i c a l I n s e r t (P/D = 2.0) , q" Versus AT. 5 * LI a a U rd - Q m = O w " h 4 a) - w x m 2 $ Momentum Exchange Model E q . (14) -- Homogeneous Model, K=l in Eq(12) Figure 63. Mercury Void Fraction as a Function of Quality for Various Flow Models. hpred. c a l c u l a t e d by Bubble Nucleat i o n Model hpred. c a l c u l a t e d by Film Evaporat ion Model P igure64 . Comparison of Heat T rans fe r Models f o r Mercury Bo i l ing a t Wetted Condit ions. hpred. c a l c u l a t e d by Bubble Nucleat ion Model hpred. c a l c u l a t e d by Film Evaporat ion Mode1 Data S c a t t e r e d Area 975OF < TSat < 1120°F x, Vapor Qual i ty F igure 65. Comparison of Heat T rans fe r Models f o r Mercury Bo i l i ng a t Wetted Condit ions. 0.50 i 158 < G < 192 Lb/Sec-Ft 0.40 0.30 0.20 Exponent a = 0.54 Determined f o r Equat ion (19) 0.15 0.10 Mercury S a t u r a t i o n P r e s s u r e , Psat, p s i F i g u r e 66. F u n c t i o n a l R e l a t i o n Be Ween hNB/ (q") 0*85 and P According t o t h e Mercury Nuc lea te B o i l i n g Data. S a t 30 5 0 70 100 200 400 600 Mercury Mass Veloc i ty , G , ~ b / ~ e c - ~ t ~ ,,O. 85 0.54)-j F igure 67. Funct iona l Re la t ion Between h 1 [(q ) (PSat NB and G According t o t h e Mercury Nucleate Boi l lng Data. lo5 Heat Flux, q", ~ t u / h r - f t 2 Figure 68. Cor re l a t ion P l o t of Mercury Nucleate Boil ing Data. 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1-0 Quality, x, Dimensionless Figure 70. Mercury Critical Heat Flux Results and Correlations Compared with Potassium Results from References (5) and (6). 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1 .0 ( 1 - x)xc Dimensionless Group, ( 1 - xc)x F i g u r e 73. C o r r e l a t i o n P l o t of T r a n s i t i o n B o i l i n g Data by Assuming Flow Model with Cons tan t D r o p l e t S ize . 0 0.1 0.2 0.3 0.4 0.5 0.6 0.7 0.8 0.9 1.0 (1 - x)xc Dimensionless Group (1 - xc)x Figure 74. Correlation Plot of Transition Boiling Data by Assuming Flow Model with Constant Density of Droplets. -0.6 - - 3,165 (i - 9 ) I - - H - - . (Consts- t Dro?let S i z e ) @ -- :Cozstz>t 3 r o q l e t Density) "-,, Local 4ccelerat ion F l e l d Cl 4- -1% ) g 9 s R F i g u r e 75 . L i q u i d Entrainment as Affected 57 Local AcceZerstlcn FieEd in the Transition Bailing 3egion, V) U rl n 3 0 V ) . 100 (A tn a) rl G 0 Predicted by Equation (55) .rl (A with 0.25 in . Centerbody Helix z E . rl Predic ted by Equation (55) n with No Centerbody Helix 2 - M 9) u a) 10 3 b rb !a ICc a) w 2 (d !4 E U (d 2i 1 0.03 0.1 1 P 0 I n s e r t Pitch-to-Diameter Ratio, P/Di, Dimensionless A Wire Coil I n s e r t 0 = 0.1428 Ref. (27) d) Hel ica l Vane I n s e r t Dcb = 0.25 Ref. (27) Figure 78. Measured Heat Transfer Parameter vs . I n s e r t Twist Ratio. 0.01 a 5 o Equat ion (53) , k C3 H e l i c a l Flow, \ = 2.6 'JY rn aJ +-I E .d rn B .rl n Lo - GaDi Reynolds Number , -, Dimensionless ub F igure 79. Superheated Mercury Vapor Heat T rans fe r Data i n t h e Wire Co i l I n s e r t Region. U n O O O u r CdFri d 0 \ O F r l r l o c n 4% rl 4 d Frr F r r F r r 0 0 0 U 0 0 0 cd m N r l m rl 4 4 E-c I-I rlrl H M - '2 P I N .rl ala u w r-i a, a 8 U C a) .rl a ld &I c3 a) 'f-! 3 rn rn a, $4 P4 .rl 'f-! d M- .a %la U-r-l (d a, rl M a d 0 8 .d s r-l U cd cd C r l U r l O a J m a h 0 r-l $ 8 " : ? L O . 5 L l 4 0 O r l a J d a, c 4 a ", - Homogeneous Model -,.,-Modified Martinelli-Nelson Model I I I I I I 0 0.2 0.4 0.6 0.8 1.0 Quality, x, Dimensionless Figure 86. Comparison of Experimental and Predicted Two-Phase Friction Pressure Drop Multiplier at T = 975'~. Sat 155 0 k = wi for TPM @ k = 1 for Homogeneous Model --- Modified Martinelli-Nelson Model Quality, x, Dimensionless Figure 87. Comparison of Experimental and Predicted Two-Phase Friction Pressure Drop Multiplier at TSat = 1060'~. Homogeneous Nodel --- Modified Martinelli-Nelson Model Quality, x, Dimensionless Figure 88. Comparison of Experimental and Predicted Two-Phase Friction Pressure Drop Multiplier at TSat = 1080°F. 141 < G < 163 lblsec-ftL 0 0.2 0.4 0.6 0.8 1.0 Quality, x, Dimensionless Figure 89. Comparison of Experimental and Predicted Two-Phase Friction Pressure Drop Multiplier at T = llOOO~. Sat --- 0 k = i p /p for L v TFM @ k = 1.0 for Homogeneous Model --- Modified Martinelli-Nelson Model VJ m PI f-? !3 e, Li a, .rl f-? a .I4 U 0 k = m f o r TPM @ k = 1 for - Homogeneous Model --- Modified Martinelli-Nelson Model Quality, x, Dimensionless Figure 90. Comparison of Experimental and Predicted Two-Phase Friction Pressure Drop Multiplier at TSat = 1 1 1 0 ~ ~ . l-i a .rl U l-i 10 rd o k = for @ k = 1 for Homogeneous Model ---Modified Martinelli-Nelson Model 0 0.2 0.4 0.6 0.8 1, Quality, x, Dimensionless Figure 91. Comparison of Experimental and Predicted Two-Phase Friction Pressure Drop Multiplier at T = 1120'~. Sat Figure 93. S h e l l Side and Composite Wall Thermal Res is tance of t h e GE S ing le Tube Boi le r . 7 0 60 50 40 T r a n s i t i o n 16 - Nucleate Boi l ing Region ------., 15 14 13 12 11 Boi le r Length, f e e t F igure 9 4 . Typica l Data f o r Local Heat Transfer Res is tance from Tes t Run No. 14. APPENDIX I CALCULATIONAL PROCEDURES APPENDIX I CAL,CUI,ATIONAL PROCEDURES A. be te rmina t ion of Conceptual Heat T rans fe r Regions A s shown schemat ica l ly i n F igure ( 9 5 ) , t h e t e s t s e c t i o n temperature p r o f i l e s f o r primary NaK flow and secondary mercury flow were recorded, from t h e measurements of 60 s h e l l thermocouples, 7 b o i l e r i n s e r t thermo- couples and 8 i n l e t and o u t l e t w e l l thermocouples, With t h e a i d of t h e s e temperature p r o f i l e s , t h e va r ious conceptual h e a t t r a n s f e r r eg ions d i s - cussed i n Chapter (IV) were determined a s fol lows: The l eng th of t h e subcooled l i q u i d hea t ing reg ion , LSC9 assumed numerical ly equal t o t h e d i s t a n c e from t h e s t a r t of t h e heated l eng th t o t he poin't of b o i l i n g incep t ion , zIB, was obtained d i r e c t l y from the ca l - c u l a t i o n of z IB ' A single-phase hea t t r a n s f e r c a l c u l a t i o n g ives : The mercury temperature a t t h e b o i l i n g incep t ion p o i n t , T ~ g = ~ ' was t a ined e i t h e r from t h e i n s e r t thermocouple measurement o r from t h e mercury s a t u r a t i o n curve by assuming t h a t T Hg I B is equal t o t h e s a t u r a t i o n tem- p e r a t u r e corresponding t o t h e measured mercury i n l e t p re s su re P Iig i " Thus, t h e l i q u i d mercury p re s su re change between b o i l e r i n l e t and b o i l i n g incept ion po in t was smal l and could be neglec ted . The average o v e r a l l NaK-to-mercury temperature d i f f e r e n c e i n t h e - subcooled hea t ing reg ion , ATsc, was c a l c u l a t e d a s fo l lows: wi th t h e NaK temperature a t b o i l i n g incep t ion , T ~ a ~ ~ ~ , given by: The o v e r a l l hea t t r a n s f e r c o e f f i c i e n t i n t h e subcooled hea t ing reg ion , - ' was c a l c u l a t e d from t h e i n d i v i d u a l NaK, composite wa l l and l i q u i d mercury hea t t r a n s f e r c o e f f i c i e n t s a s fol lows: ( s e e F igure 9) The s h e l l - s i d e NaK flow hea t t r a n s f e r c o e f f i c i e n t , hs, was ca l cu la t ed from Dwyer's equat ion f o r concen t r i c a n n u l i as fol lows: (33) where a = 4.82 + 0.69(Y) i nne r r a d i u s Y = o u t e r r a d i u s and Values of are provided as a function of the Reynolds number in max Reference (7) for the condition of fully developed turbulent flow of liquid metal. In the present experiment, the shell-side NaK flows were held almost constant at about 7075 lb/hr (2 3%) and the NaK average temperatures were kept at about 1270°F (+ - 3.5%). In this manner the shell-side heat transfer coefficient calculated from Equation (A5) had almost a constant value of about 2830 ~tu/hr-ft~-'~ (+- 1%). In Figure ( 9 4 ) , the shell- side and composite wall thermal resistances of the single test boiler are given. A shell flow and wall combined heat transfer coefficient, U was then calculated to have a value of 1520 ~tu/hr-ft~-'~ when CW referred to the tantalum boiler tube inside diameter. Lyon's equation was employed to calculate the mercury liquid heat transfer coefficient, h ~ g ~ as follows: where GH, denoted as the helical mass velocity is: and the equivalent diameter, De, following its usual definition, is evaluated as : Equations (Al) through (A6) were then used t o p r e d i c t t h e lengt l l of t h e subcooled hea t ing r eg ion , L . S C The c r i t i c a l tieat f l u x p o i n t , z a s shown i n F igure (95) , was c ' determined by assuming t h a t t h e maximum hea t f l u x occurs a t t h i s p o i n t . This was done by employing a l o c a l g r a p h i c a l d i f f e r e n t i a t i o n along t h e smooth curve of t h e NaK temperature p r o f i l e , s i n c e t h e l o c a l hea t f l u x is p ropor t i ona l t o t h e s h e l l - s i d e flow temperature g r a d i e n t . The a x i a l p o s i t i o n a t which vapor superhea t ing commences, z IS is assumed t o be t h e po in t a t which t h e mercury temperature begins t o r i s e , a s shown i n F igure (95). F i n a l l y , t h e l e n g t h s of t h e v a r i o u s hea t t r a n s f e r r eg ions of a once-through b o i l e r a r e determined a s fo l lows: Length of subcooled l i q u i d r eg ion L = z s c I B Length of n u c l e a t e b o i l i n g r eg ion = z - z L~~ c IB Length of t r a n s i t i o n b o i l i n g r eg ion LTB = z - z I S C Length of superheated vapor r eg ion LSN = z - z t I S With t h e t o t a l b o i l e r heated l e n g t h z t = 16 f e e t and predetermined va lues of L and L f o r each t e s t run, t h e nuc l ea t e b o i l i n g and sc SH t r a n s i t i o n b o i l i n g r eg ions were e a s i l y determined by: LNB = zc - L and L T B = 1 6 - s c Lsc - L~~ - L~~ B. Average and Overa l l Resu l t s - AVERAGE HEAT FLUX, q" The n e t r a t e of hea t t r a n s f e r f o r t h e mercury b o i l e r w a s given = n e t 'N~K'N~K (Al l ) where QLoss, t h e hea t l o s s , w a s obtained a t t h e average NaK temperature from t h e hea t l o s s c a l i b r a t i o n runs given i n F igure (19). The hea t t r a n s f e r r e d i n t h e subcooled l i q u i d reg ion w a s g iven by: = W C Qsc HgHg T ~ g ~ ~ - T ~ g i (A12 1 where T HgIB was eva lua ted from t h e s a t u r a t i o n curve by knowing P Ngi" The amount of t h e hea t t r a n s f e r r e d i n t h e superheated vapor reg ion , QSH, was determined from t h e mercury temperature r i s e a s follows: = W C Hg Hgv (A13 1 where T ~ g l ~ was provided by t h e mercury temperature p r o f i l e as measured by t h e b o i l e r i n s e r t thermocouples. Ilence, t h e n e t hea t t r a n s f e r r e d t o b o i l t h e mercury w a s calcuJtated With t h e hea t ing l e n g t h s f o r t h e i n d i v i d u a l h e a t t r a n s f e r r eg ions de- termined, t h e average hea t f l u x i n each reg ion was t h u s c a l c u l a t e d by t h e s e equat ions : 171 - - Qsc '"sc PDiLsC , wi th LB = z t - Lsc - L s ~ ~ Furthermore, t h e average h e a t f l u x i n t h e n u c l e a t e b o i l i n g r eg ion was obtained from t h e NaK temperature p r o f i l e provided by t h e b o i l e r s h e l l thermocouple measurements, Assuming t h a t t h e temperature g r a d i e n t a long t h e s h e l l i s approximately equa l t o t h e s h e l l - s i d e f lowing N a K temperature g r a d i e n t , t h e fo l lowing average s h e l l temperature g r a d i e n t was used t o c a l c u l a t e t h e average hea t f l u x i n t h e n u c l e a t e b o i l i n g reg ion : where t h e average hea t f l u x due t o hea t l o s s e s i s g iven by: Hence, w i th t h e hea t ing l e n g t h of t h e n u c l e a t e b o i l i n g r eg ion determined, t h e amount of hea t t r a n s f e r r e d i n t h e n u c l e a t e b o i l i n g r eg ion was e a s i l y ca l cu l a t ed by: F i n a l l y , t h e hea t t r a n s f e r r e d i n t h e t r a n s i t i o n b o i l i n g r eg ion was ob- t a ined from t h e fol lowing equa t ion : and the average heat f l u x i n t h i s region was obtained by: with LTB calcula ted from Equation (A10). AVERAGE OVERALL HEAT TRANSFER COEFFICIENTS, With t h e average heat f luxes f o r each individual heat t r a n s f e r - region determined, t h e average o v e r a l l hea t t r a n s f e r c o e f f i c i e n t s , U, f o r each region were ca lcula ted by: - = 78 / E S C osc where Eosc is t h e average o v e r a l l NaK t o mercury temperature d i f fe rence i n the subcooled l i q u i d region ca lcula ted by: The temperature no ta t ion i n the above equation is i l l u s t r a t e d i n Figure ( 9 5 ) . Similar ly , t h e average o v e r a l l hea t t r a n s f e r c o e f f i c i e n t s for t h e nuclea te bo i l ing , t r a n s i t i o n bo i l ing , and superheated vapor regions were ca lcula ted a s follows: The average o v e r a l l NaK t o mercury temperature d i f f e r e n c e , ( 2 , was obtained from t h e fo l lowing equa t ions : where t h e NaK temperature a t t h e p o i n t where vapor superhea t ing s t a r t s , T ~ a ~ ~ ~ ' and t h e N a K temperature a t t h e c r i t i c a l hea t f l u x p o i n t , T ~ a ~ c " a r e : and wi th QSH and QTB p r ev ious ly known from Equations (Al3) and (A21), r e s p e c t i v e l y . AVERAGE 79ERCUP.Y HEAT TRANSFER COEFFICIENTS. The average mercury hea t t r a n s f e r c o e f f i c i e n t f o r each hea t t r a n s f e r reg ion was c a l c u l a t e d from t h e average o v e r a l l hea t t r a n s f e r c a e f f i c i e n . t , - - U, and t h e combined wa l l and s h e l l - s i d e hea t t r a n s f e r c o e f f i c i e n t , Ucw' For in s t ance , i n t h e subcooled l i q u i d r eg ion was ca l cu la t ed by t h e S C equat ion : - where t h e combined w a l l and s h e l l - s i d e hea t t r a n s f e r c o e f f i c i e n t , uCW' is given by: S imi l a r ly , t h e average mercury hea t t r a n s f e r c o e f f i c i e n t s i n re- gions o the r than t h e subcooled l i q u i d reg ion were then s u c c e s s f u l l y ca l cu la t ed a s fol lows: - - with t h e predetermined va lues of EHb9 UTB, and USH g iven by Equations (A25), (A26), and (A27) r e spec t ive ly . OVERALL PRESSURE DROP, (AP)o Two Taylor s l a c k diaphragm p res su re t r ansduce r s were provided f o r measuring t h e o v e r a l l p r e s su re drop a c r o s s t h e t e s t s e c t i o n . The seven b o i l e r i n s e r t thermocouples employed f o r measurement of mercury tempera- t u r e i n t h e b o i l i n g r eg ion can a l s o be used t o e s t i m a t e t h e two-phase b o i l i n g p re s su re drop wi th t h e a i d of t h e mercury s a t u r a t i o n curve. The o v e r a l l p r e s s u r e drop a c r o s s t h e test s e c t i o n can g e n e r a l l y be expressed a s fo l lows: (APb = (APITp + (APISpV * (Ap)SpL + (AP) eleve + (ApImiSc (A381 where (AP)Tp = Pres su re l o s s due t o two-phase b o i l i n g . SPV = Pres su re l o s s due t o s ingle-phase vapor. SPL = Pres su re l o s s due t o s ingle-phase l i q u i d . ( A P ) e ~ e v , = Eleva t ion l o s s f o r bo th s ingle-phase l i q u i d and vapor , and two-phase f l u i d . (AP)misc. = Miscel laneous l o s s e s such a s j o i n t s , c o n t r a c t i o n , expansion, e t c . , f o r both s ingle-phase l i q u i d and vapor. (Ap) = Overa l l p r e s su re drop measured by Taylor p r e s s u r e t r ansduce r s . By c a r e f u l l y examining t h e running p ipe c o n s t r u c t i o n s and geometr ies between t h e two Taylor p r e s s u r e t r ansduce r s , t h e misce l laneous p re s su re drops , i . e . , j o i n t s , c o n t r a c t i o n , and expansion, were es t imated by t h e convent iona l p r e s su re drop formulas and found t o be i n t h e range of 1 .21 t o 2.28 p s i f o r v a r i o u s mercury flow r a t e s . The test s e c t i o n and p ipe layouts of t h e p re sen t experiment f e a t u r e d a 7 degree s l o p e o r 5/8-inch per f o o t e l e v a t i o n between t h e two Taylor t r ansduce r s . Therefore , t h e e l e v a t i o n l o s s was l a r g e , e s p e c i a l l y f o r l i q u i d mercury, and had t o be taken i n t o account. For v a r i o u s mercury flow r a t e s , ("1 e l ev . was es.- t imated t o be i n t h e range of 7 t o 8.7 p s i f o r l i q u i d mercury and 0.5 t o 1.56 p s i f o r mercury vapor . 176 Tn t h e t e s t s e c t i o n , t h e f r i c t i o n a l p re s su re arul s t a t i c head l o s s e s of t h e subcooled licluid were n e g l i g i b l y small due t o a very s h o r t sub- cooled l i q u i d reg ion f o r a l l t e s t runs ( i n t h e . r a n g e of 0 .4 t o 1 f o o t ) and a very small ang le of test s e c t i o n o r i e n t a t i o n (7 degrees from h o r i z o n t a l p o s i t i o n ) . On t h e con t r a ry , t h e vapor p re s su re l o s s i n t h e s ingle-phase superheated vapor reg ion c o n s t i t u t e d a l a r g e sha re of t h e t o t a l b o i l e r p re s su re drop. For most t e s t runs , t h e superheated vapor reg ion covered a l a r g e po r t ion of t h e t e s t b o i l e r , gene ra l ly occupying both t h e h e l i c a l i n s e r t reg ion and t h e wire c o i l region. The super- heated vapor p re s su re l o s s can be gene ra l ly expressed by i ts momentum and f r i c t i o n a l components a s fol lows: (APISPV = (AP)SPVm + [(Ap)sPvf h e l i c a l i n s e r t reg ion (A401 wire c o i l reg ion The term (AP) SPVm, which accounts f o r t h e momentum p res su re l o s s of mercury vapor from t h e superheated reg ion t o t h e b o i l e r e x i t , was ca l cu la t ed by t h e fol lowing equat ion: where G def ined a s t h e a x i a l mass v e l o c i t y , is t h e quo t i en t of mercury a " mass flow rate and t h e n e t flow a rea . For t h e b o i l e r tube conta in ing a h e l i c a l i n s e r t , t h e n e t flow a r e a , AF, i s given i n terms of t h e i n s e r t centerbody d iameter , DcbD and t h e i n s e r t t ape th i ckness , A t , a s fo l lows: P , def ined a s t h e mean square r a t i o of h e l i c a l v e l o c i t y t o a x i a l v e l o c i t y , was used i n Equation (A41) as an empi r i ca l f a c t o r t o account f o r t h e e f f e c t of t h e h e l i c a l flow condi t ion on momentum f l u x eva lua t ion . For any c r o s s s e c t i o n along t h e b o i l e r tube , t h e t o t a l momentum is given by a combination of l i n e a r momentum and angular momentum, The v e l o c i t y i n t h e h e l i c a l flow i s considered s i m i l a r t o t h a t of a s o l i d c y l i n d r i c a l body r o t a t i n g about i t s a x i s and t r a n s l a t i n g a x i a l l y i n t h e d i r e c t i o n of flow i n such a way t h a t t h e r e s u l t a n t v e l o c i t y vec to r a t any r a d i u s i s p a r a l l e l t o t h e h e l i c a l vane. Therefore, t h e l i n e a r momentum is independent of tube r a d i u s and h e l i c a l angle , whereas t h e angular momentum is dependent, Thus, a mathematical express ion f o r t h e t o t a l momentum f l u x a c r o s s any c r o s s s e c t i o n can be obtained by a sur- f a c e i n t e g r a t i o n of l o c a l momentum f l u x over t h a t c o n t r o l su r f ace . To avoid any mathematical d i f f i c u l t y i n eva lua t ing such a s u r f a c e i n t e g r a l , an empir ica l f a c t o r was thus introduced t o o b t a i n t h e mean t o t a l momentum f l u x over any c r o s s s e c t i o n a l a r ea . We d e f i n e t h a t : (AP)m = t o t a l momengurn p re s su re l o s s L given by t h e fol lowing r e l a t i o n : For a s o l i d body r o t a t i o n , t h e h e l i c a l v e l o c i t y , VH, can be w r i t t e n a s : and t h e h e l i c a l c r o s s s e c t i o n a l a r e a , dAHs can be obtained from t h e v a mass flow con t inua t ion a s d (2lIr)dr. S u b s t i t u t i n g V and d q l H i n t o equat ion (A44) and i n t e g r a t i n g , g ives : where D is t h e diameter of t h e centerbody of t h e b o i l e r i n s e r t . cb The f r i c t i o n a l p re s su re l o s s f o r mercury vapor i n t h e h e l i c a l vane i n s e r t reg ion can be est imated by us ing t h e convent ional p re s su re drop c o r r e l a t i o n f o r tube and mul t ip ly ing by a c o e f f i c i e n t t o account f o r t h e h e l i c a l i n s e r t e f f e c t . This approach was s u c c e s s f u l l y used t o c o r r e l a t e t h e p re s su re drop d a t a f o r water (29) and a i r ( 2 7 1 flows. The convent ional Fanning type p re s su re c o r r e l a t i o n g ives : with f e = 0.316 14 (equiva len t f r i c t i o n f a c t o r ) (A47) (helical mass velocity) LH = L (helical path length) Alternatively, Equation (A46) can be rewritten as: with (AP)a = (plain tube) K is then defined as a pressure loss multiplier to account for the in- P crease in pressure loss due to helical flow conditions. The frictional pressure loss for mercury vapor in the wire coil insert region is complicated, and as yet no analytical prediction or correlation is available. Furthermore, it was found in Reference (27) that the use of a helical flow pressure multiplier to correlate the pressure drop for air flow with wire coil inserts resulted in an under- estimation of the test data, The only way to predict the frictional loss for wire coil insert is to run an experiment (such as with heated air) to obtain the actual value of the multiplier. A series of hydraulic tests using water and heated air with various geometries of wire coil inserts was completed recently, (27)3(29) and some of the test results a r e shown i n F igure (96). A s shown i n F igure (97) , t h e exper imenta l ly determined p re s su re l o s s m u l t i p l i e r , K was found t o be approximately P ' 22 f o r t h e p re sen t wire c o i l geometry, and was found n o t t o be a s t r o n g func t ion of t h e Reynolds number. Thus, t h e mercury vapor p r e s s u r e l o s s due t o f r i c t i o n i n t h e w i r e c o i l r eg ion was c a l c u l a t e d by t h e fo l lowing empi r i ca l equat ion: (AP) s,vf = 22 (AP) a dw ( f o r P/D = 0.97 and - = 0.1428 wire c o i l i n s e r t ) Di Wit11 t h e va r ious p r e s s u r e l o s s e s , (AP) SpV, (AP) (AP) eleve and (AP)misc. , determined by t h e above equa t ions , t h e p re s su re l o s s i n t h e two-phase b o i l i n g r eg ion w a s es t imated by s u b t r a c t i n g t h e s e known pres- s u r e l o s s e s from t h e t o t a l measured p r e s s u r e l o s s which was recorded by two Taylor t ransducers . For comparison, t h e two-phase p re s su re l o s s w a s a l s o ob ta ined from t h e mercury s a t u r a t i o n curve by using t h e i n s e r t thermocouple temperature read ings i n t h e b o i l i n g reg ion . The mercury s i n g l e phase vapor p re s su re l o s s due t o f r i c t i o n , ( P P ~ ~ ~ ~ , was es t imated from Equation (A38) w i th t h e assumptions (1) t h a t t h e t.wo- phase p re s su re l o s s va lues ob ta ined from i n s e r t thermocouples were c o r r e c t , and (2) t h a t t h e o t h e r p r e s s u r e l o s s e s involved were eva lua ted w i th r e l i a b l e equa t ions . S ince t h e p re s su re l o s s m u l t i p l i e r , K , was assumed P t o account f o r p r e s su re l o s s i n t h e h e l i c a l vane i n s e r t reg ion , t h e f r i c t i o n a l p r e s su re drop of mercury vapor i n t h e w i r e c o i l r eg ion was p red i c t ed from Equation ( A 4 0 ) . With t h i s p red ic ted va lue , t h e p re s su re l o s s m u l t i p l i e r (def ined a s K = (AP)wc/(AP)a) was then c a l c u l a t e d and P compared w i t h e a r l y a i r h e a t i n g test d a t a f o r t h e same w i r e c o i l i n s e r t . The two-phase b o i l i n g p r e s s u r e d r o p , (bP)Tp, c o n s i s t s of two com- ponen ts , i . e . , f r i c t i o n a l and momentum, and i s r e p r e s e n t e d by: The momentum t e r m , (AP)Tp,) was c a l c u l a t e d by i i l t e g r a t i n g the d i f - f e r e n t i a l form of momentum p r e s s u r e d rop between t h e q u a l i t y x = 0 and x = x t o y i e l d : e w i t h K d e f i n e d as t h e s 3 i p r a t i o , i .e . , t h e average v e l o c i t y r a t i o of vapor t o l i q u i d . For a l l t h e t e s t r u n s , t h e mercury was always superheatecl 20Q°F t o 400°F a t t h e b o i l e r e x i t . T h e r e f o r e xe was always e q u a l t o one and 1 1 Equat ion (A56) reduced t o s imply - = -. Equa t ion (A55) was t h e n 'e 'VIS r e w r i t t e n as: Subsequent ly , t h e f r i c t i o n a l two-phase p r e s s u r e d rop , (AP)Tpf, was c a l c u l a t e d from Equa t ion ( A 5 4 ) , and a n i n t e g r a t e d two-phase f r i c t i o n - p r e s s u r e d rop m u l t i p l i e r , , w a s e v a l u a t e d by t h e f o l l o w i n g e q u a t i o n : exp where (AP) denotes t h e s ingle-phase l i q u i d p re s su re drop which would L r e s u l t i f t h e f low r a t e of t h e l i q u i d i n t h e tube were equa l t o t h e a c t u a l two-phase mix ture f low r a t e , and is expressed by: -1/4 w i th f L = 0,316 (A601 and LB = LNB + LTB = z z I S - I B The exper imenta l ly ob ta ined two-phase f r i c t i o n p re s su re drop mul t i - p l i e r s were then compared wit11 t h e v a l u e s from t h e M a r t i n e l l i model ( 3 4 ) - modified f o r mercury, % M , and t h e va lues from a homogeneous flow model, - - 6 Mathematically and 2 li were obta ined by a numerical i n t e g r a t i o n - H. from t h e i r l o c a l va lues , Q and QH, a t s e l e c t e d temperatures on t h e Q M ve r sus x p l o t s which are g iven i n F igures (102) and (104). The d e t a i l e d d i s cus s ion about t h e s e l o c a l f u n c t i o n s QN (T,x) and QH (T,x) i s given i n Sect i o n (D) . - For t h e p re sen t one hundred percent b o i l i n g runs , $ and were c a l c u l a t e d by t h e fo l lowing equa t ions : and t h e r e s u l t s a r e p l o t t e d i n F igures (98 and (99). C. Local Heat Transfer Resu l t s By us ing t h e s h e l l temperature p r o f i l e s and assuming t h a t t h e - d T ~ a ~ a x i a l bulk NaR temperature g r a d i e n t , - dz , is equal t o t h e a x i a l dTs s h e l l temperature g r a d i e n t , -dz , t h e l o c a l hea t f l u x q" was ca l cu la t ed by t h e fol lowing equat ion: where q" is t h e hea t f l u x obtained by d iv id ing t h e t o t a l hea t Posses Loss of t h e t e s t s e c t i o n by t h e a c t i v e hea t t r a n s f e r a r ea . The va lues of d T S - dz needed f o r computation of q'?n Equation (A64) were obtained by g raph ica l d i f f e r e n t i a t i o n of a smooth curve drawn through t h e s h e l l temperature po in t s . The l o c a l v a r i a t i o n of hea t f l u x i n t h e subcooled hea t ing reg ion could not be obtained d i r e c t l y s i n c e t h e l eng th of t h i s reg ion was too s h o r t t o permit accu ra t e measurements. However, t h e hea t f l u x a t t h e mercury i n l e t was est imated by t h e fol lowing equat ion: where 5 t h e average o v e r a l l hea t t r a n s f e r c o e f f i c i e n t , was g iven by s c ' Equation ( A 2 3 ) , and TNaK and THgi were provided by w e l l tllermocouple readings. The remainder of t h e l o c a l h e a t f l u x p r o f i l e i n t h e subcooled hea t ing - reg ion was est imated through t h e u s e of average h e a t f l u x , q'' i n t h e s c reg ion given by Equation (A15), The a x i a l p o s i t i o n a t which t h e c r i t i c a l hea t f l u x cond i t i on occurred and t r a n s i t i o n b o i l i n g commenced, z was c ' determined from t h e s h e l l temperature p r o f i l e a s t h a t po in t a t which a marked change occurred i n t h e l o c a l s h e l l temperature g rad ien t . From t h e many s h e l l thermocouple readings along t h e t e s t s e c t i o n i t was ob- dTs served t h a t t h e l o c a l va lues of -d z increased cont inuously from t h e b o i l i n g incep t ion p o i n t , z up t o i t s maximum value and then ab rup t ly IB dTs decreased, The po in t a t which -dz reached i t s maximum va lue marked t h e c r i t i c a l hea t f l u x cond i t i on , and t h e c r i t i c a l h e a t f l u x was then de- termined by Equation ( A G 4 ) . I Equation ( A 6 4 ) is u s e f u l only up th rough the t r a n s i t i o n reg ion be- cause i n t h e superheated reg ion t h e s h e l l temperature p r o f i l e becomes dT S i nc reas ing ly f l a t , and t h e l o c a l d i f f e r e n t i a t i o n of -dz becomes very d i f f i c u l t and inaccura te . Therefore, i n t h e superheated reg ion t h e l o c a l hea t f l u x was determined by t h e fol lowing equat ion wi th t h e he lp of t h e mercury temperature p r o f i l e s : The l o c a l bulk WaK temperature was obtained from t h e known bulk i n l e t temperature, T ~ a ~ i (measured by we l l thermocouple), and t h e in- t e g r a t i o n of t h e l o c a l hea t f l u x wi th r e spec t t o z , a s fol lows: -#=5 S O O R 3s t h e a x i a l d i s t r i b u t i o n of q" was c b t a l x c d ss d t s c r i b e d ,ibove, :he i n t e g r a t i o n e x p r e s s e d i n E q u a t i o n ( A 6 7 ) T--as c a l c u l a t e d n u m e r i c a l l y 25 f o l l o v s : T:-.e Loca l q u a l i t y , x, L-2s c a l c u l a t e d by a c o c s i d e r a t i c n o f e n e r g y b a l a n c e 2 s f o l l o w s : Tor a l l s u p e r h e a t e d v a p c r e x i t c o n d i t i o n s i n rke p r e s e n t exper i r r t sn ts , s was t a k e n a s the q t i a l i t g zit t h e b e g i n n i n g cf f l u i d b u l k supeshi?atiiag C (2 IS ) where x was a s s u z e d t o b e 100. Hence, Equation (A69) r educed t o : 0 s a t u r a t i o n tempera tures , and t h e h e l i c a l v e l o c i t i e s , V and VHIS, a r e H eva lua ted by t h e fol lowing equa t ions : Local va lues of t h e o v e r a l l hea t t r a n s f e r c o e f f i c i e n t , U , were c a l c u l a t e d from t h e l o c a l va lues of q" and AT, a s fol lows: where AT is obta ined from t h e c a l c u l a t e d bu lk NaK temperatures and t h e mercury temperatures measured by i n s e r t thermocouples. F i n a l l y , t h e l o c a l v a l u e s of mercury hea t t r a n s f e r c o e f f i c i e n t , h were c a l c u l a t e d by s t ib t rac t ing t h e thermal r e s i s t a n c e of t h e flow- Hg ' ing NaK, s t a t i c NaK l a y e r , s t a i n l e s s steel annulus , and tantalum tube wal l from t h e o v e r a l l thermal r e s i s t a n c e , a s fo l lows: where t h e combined thermal r e s i s t a n c e , - , was g iven i n equa t ion ( A 3 4 ) . - ucw D. Local Two-Phase P re s su re Drop Resu l t s The l o c a l va lue of s a t u r a t i o n p re s su re dur ing two-phase b o i l i n g of mercury was obtained by us ing measured mercury s a t u r a t i o n tempera tures t oge the r wi th t h e vapor p re s su re curve c i t e d from Reference (12) . The 187 l o c a l b o i l i n g p re s su re g rad ien t w a s c a l c u l a t e d from t h e l o c a l TP b o i l i n g temperature g r a d i e n t and t h e vapor p re s su re curve f o r mercury. In terms of t h e d e r i v a t i v e of temperature w i th r e spec t t o p re s su re along t h e vapor p re s su re curve, t h e l o c a l b o i l i n g p re s su re g r a d i e n t was ob- ta ined a s fol lows: T~~ obtained by g r a p h i c a l d i f f e r e n t i a t i o n along with 2 and t h e mercury temperature p r o f i l e and t h e mercury vapor p re s su re curve, r e spec t ive ly . A knowledge of t h e l o c a l momentum and e l e v a t i o n p re s su re g r a d i e n t s is needed t o c a l c u l a t e l o c a l va lues of t h e two-phase f r i c t i o n a l p re s su re g rad ien t from t h e t o t a l two-phase p re s su re g r a d i e n t , a s fol lows: Considering an annular flow model f o r t h e b o i l i n g reg ion , i . e . , l i q u i d flowing i n an annulus surrounding a c o r e of vapor , t h e momentum p res su re g rad ien t f o r two-phase mixture w a s def ined: and t h e e l e v a t i o n p re s su re g rad ien t was def ined: = - E L s i n 0 pVa + pL ( 1 - a) c where 8 , t h e e l e v a t i o n a n g l e of t h e test s e c t i o n , is 7 degrees i n this experiment . The void fraction,^, i n equa t ions (A77) and (A78) was e l imina ted by employing t h e c o n t i n u i t y equa t ion of two-phase mix tures a s fo l lows: By us ing t h i s express ion f o r a , Equations (A77) and (A7S) w e r e r e w r i t t e n as: and 01 - g s i n 8 x + K (1-x) gc x + K (1-x) " v By assuming K , t h e s l i p r a t i o , cons t an t f o r a p a r t i c u l a r run, t h e r i g h t hand s i d e of Equation (A81) was d i f f e r e n t i a t e d y i e l d i n g : The proper v a l u e of K is u n c e r t a i n a t p r e sen t . Some i n v e s t i g a t o r s (5), (35) found t h a t K = (pV/PL)1'2 i n t h e i r ana lyses . A homogeneous model, assuming simply K = 1, is u s u a l l y used i n t h e p r e d i c t i o n of two-phase 189 pres su re drops. For a homogeneous model, Equations (A81) and (A82) were s imp l i f i ed to : and i n Equation ( A 8 4 ) , t h e l o c a l q u a l i t y v a r i a - I n order t o c a l c u l a t e t i o n must be known. A cons ide ra t ion of energy ba lance g ives : * With dZ d T ~ ~ ----Aevalua ted l o c a l l y by g r a p h i c a l d i f f e r e n t i a t i o n and dz from t h e i r p r o f i l e s presented i n F igures (41) t o (56) and Figures (21) t o ( 3 6 ) , r e spec t ive ly . The b o i l i n g l eng th (L + L ) was g iven B = L~~ TB previous ly i n Equation (A10) f o r each experimental run. The mass v e l o c i t y , 6 , used i n Equation (A84) must be co r r ec t ed t o account f o r t h e h e l i c a l flow e f f e c t . A s d i scussed previous ly , an equat ion was proposed f o r G a s fol lows: given i n Equation (A45). Then, t h e f r i c t i o n a l p re s su re was c a l c u l a t e d from Equation (A76). The l o c a l two- phase p re s su re drop m u l t i p l i e r is t h e r a t i o of two-phase t o s ingle-phase l i q u i d f r i c t i o n a l p re s su re g rad ien t : where the liquid frictional pressure gradient is calculated by the con- ventional Fanning equation modified for helical flow conditions as follows: with One of the earliest efforts in the prediction of two-phase pressure drop was that of Martinelli et al. (37) in which a substantial amount of two- phase, two-component data was successfully correlated. A further im- provement of this work by Lockhart-Martinelli (37) gave the most widely applied method for predicting two-phase pressure drop. Later, bfartinelli- Nelson ( 3 4 ) successfully correlated their forced-convection boiling water data with the method developed in Reference (37). They then constructed a series of curves by plotting the two-phase multiplier defined in Equation (A88) versus the quality with the saturation pressure (or temperature) used as a parameter. The Martinelli-Nelson curves, shown in Figure (loo), were prharily deduced from high pressure steam data; consequently, these curves were proposed for use only with steam. However, these curves can be generalized and used for many other fluids including liquid metals. In References (5) and ( 6 ) , the Martinelli-Nelson curves were modified successfully to correlate the potassium boiling pressure drop data. An attempt was then made to modify the Martinelli-Nelson curves for predicting the current mercury boiling pressure drop data. This modification is described below. I 4 Introducing a scale factor, defined , the Martinelli- Nelson curves were replotted in Figure (101) to show the two-phase multiplier (4 ) as a function of the scale factor, with the quality used as a parameter. The scale factor, which involves only the physical properties for both phases of any single- or two-component, two-phase fluid, was used as a common basis to scale out the two-phase multiplier for a particular fluid in a forced-convection boiling process. Since the scale factor is primarily temperature (or pressure) dependent, there are corresponding mercury temperatures for selected values of the scale factor, as shown in Figure (101). Based upon this transformation, a series of curves for mercury similar to the Martinelli-Nelson curves for water were constructed in Figure (102). Another frequently used two-phase pressure drop correlation is the so-called "homogeneous model" which was derived from the definition of the two-phase pressure drop multiplier as follows: In Equation (A91) the Fanning type pressure drop correlation and the Blasius relation for the friction factor have been used. With the density and viscosity for a homogeneous two-phase mixture written as: 192 and Equation (A91) becomes: Equation (A94) is a general expression for predicting two-phase, pres- sure drop for any single- or two-component, two-phase fluid. In Figure (103) Equation (A94) was generalized by using the scale factor. A plot similar to that of the Martinelli-Nelson curves is given in Figure (104) for various mercury temperatures. Vapor Subcooled Liquid T r a n s i t i on T ~ a ~ i T ~ a ~ ~ s T ~ a ~ c T Hgo . -* -. - - - T ~ g I s O Lsc 'IB L~ z c L~~ z IS L~~ 16 Dis tance Along B o i l e r Tube from Mercury I n l e t , z , f t Figure 95. Diagrammatic D e f i n i t i o n of Terminology Employed i n t h e Ca lcu la t iona l Procedures f o r Counter-Current Once-Through Mercury Boi le r . 7 4 AWire Co i l P/Di = 0.83, No Centerbody helix P/Di = Le46 0.25 Centerbody 2 8 P l a i n Tube Equation (A47) (Blas ius Cor re l a t ion ) Axial Flow Reynolds No., GaDi , Dimensionless 1.1 Figure 96. F r i c t i o n Fac to r V s . Axial Flow Reynolds Number f o r Tubes Containing I n s e r t s wi th Forced A i r Flow. 0.03 0 .1 1 10 Insert Pitch-to-Diameter Ratio, P / D ~ , Dimensionless 0 A Wire Coil Insert 0 U! = 0.1428 Ref. (27) @ Helical Vane Insert Dc, = 0.25" Ref. (27) Figure 97. Measured Pressure Drop Parameter vs. Insert Twist Ratio, Q u a l i t y , x , % Figure 100. Rat io of Two-Phase F r i c t i o n P res su re Gradient t o Liquid-Phase Gradient a t t h e Same Mass Veloc i ty f o r Water a t Various P re s su res and Q u a l i t i e s . (Reference 34) Water Temp. (OF) Figure 101. Generalized Martinelli-Nelson Model Two-Phase Frictional Pressure Drop Multiplier. %, Scale Factor, , Dimensionless Figure 101 (Cont'd). Generalized Martinelli-Nelson Model Two-Phase Frictional Pressure Drop Multiplier for the Region of 0.5 < x 1.0, Figure 102. Two-Phase Frictional Pressure Drop Multiplier for Mercury from Modified Martinelli-Nelson Model. Figure 103. Generalized Homogeneous Model Two-Phase Frictional Pressure Drop Multiplier. Quality, x, Dimensionless Figure 104. Two-Phase Frictional Pressure Drop Multiplier for Mercury from Homogeneous Model (n = 0.25). APPENDIX I1 TABULATION OF EXPERIMENTAL DATA TABLE 1 LOOP THERMOCOUPLE DESCRIPTION - - Thermo- Wire couple Junc t ion Outs ide S i ze Number Type Dia. , i n . (GA) 1 CNG 0.062 28 2 CNG 0.062 28 3 CNG 0.062 28 4 CNG 0.062 28 5 CNG 0.062 28 6 CNG 0.062 28 7 CNG 0.062 28 8 CNG 0.062 28 9 CNG 0.062 28 4 9 OSJ 0.125 24 5 0 OSJ 0.125 24 5 1 OSJ 0.125 24 52 OSJ 0.125 2 4 5 3 . OSJ 0.125 24 54 OSJ 0.125 2 4 5 5 OSJ 0.125 24 5 6 OSJ 0.125 24 60 OLJ See Note (2) 24 6 1 OLJ See Note (2) 24 101 CNG 0.062 2 8 102 CNG 0.062 2 8 103 CNG 0.062 2 8 104 CNG 0.062 28 106 OSJ 0.125 24 116 OSJ 0.125 24 117 OSJ 0.125 24 118 OSJ 0.125 24 119 OSJ 0.125 24 120 OSJ 0.125 24 124 OSJ 0,125 24 D i g i t a l Channel Locat ion B o i l e r I n l e t Well Bo i l e r I n l e t Well Bo i l e r E x i t Well B o i l e r Ex i t Well Condenser Upper Drum Well Condenser Lower Drum Well Pump E x i t Well P rehea t e r I n l e t Well P rehea t e r Ex i t Well P re s su re Gage P6 (Pump E x i t ) P re s su re Gage P7 (Liq. DifE. High) P re s su re Gage P7 (Liq. Dif f . Low) P r e s s u r e Gage P8 (Boi le r I n l e t ) P re s su re Gage P3 (Boi le r E x i t ) P re s su re Gage P4 (Vap. D i f f . High) P r e s s u r e Gage P4 (Vap. D i f f . Low) P re s su re Gage P5 (Cond. Top Drum) Reference Box B Reference Box C Bo i l e r I n l e t Bo i l e r I n l e t Bo i l e r E x i t Bo i l e r E x i t Gas Heater I n l e t P re s su re Gage P1 (Heater I n l e t ) P re s su re Gage P2 (Heater D i f f . High) P re s su re Gage P2 (Heater D i f f . Low) Flowmeter Duct Flowmeter Magnet S t a t i c NaK Expansion Tank ( l )Heat Junc t ion Reference Temperature, 1 5 0 ' ~ f o r a l l T/CVs (2)0. 125-inch OD Cu Wire (Al l o t h e r a r e Chromel-Alumel) ( 3 ) ~ u n c t i o n Type: CNG Capped and Non-Grounded OSJ Open Surface Junc t ion OLJ Open Loop Junc t ion TABLE 2 BOILER INSERT THERMOCOUPLE DESCRIPTION * S u r f a c e Spot Weld 208 TABLE 3 BOILER INSERT THEm4OCOUPLE DESCRIPTION -- *Capped N o t G r o u n d e d TABLE 4 *See Nomenclature for symbols used in the data table. Units given in Tables 4 to 12 are given as follows: AR Radial Acceleration g ' s G Mass Velocity lbmlsec-it2 h,; Heat Transfer Coefficient ~tu/hr-ft'-~~ L Length ft P Pressure psia q" Heat Flux Btu/hr-f t2 Cj Heat Transfer Rate kilo-watts T Temperature OF TABLE 5 OVERALL PRESSURE DROP RESULTS *Calcula ted from t h e i n l e t and e x i t Taylor p r e s s u r e t r a n s d u c e r s **Estimated from measured s a t u r a t i o n temperatures TABLE 6 OVERALL PRESSURE DROP RESULTS IN TWO PHASE REGION Run * NO. ~g T ~ a t 'sat (A?),, ( A F ) ~ ~ m (WL ( A P ) ~ ~ f (I' exp "Obtained from insert thermocouple readings. TABLE 7 SUBCOOLED L I Q U I D RESULTS G T AT h~ N~~ N I; Run No. Date T i m e Hg Sat s c Pe s c TABLE 8 LOCAL NUCLEATE BOILING HEAT TRANSFER RESULTS Run T ~ a t G L~~ qlt X NO. ~g h~~ ( T ~ T ~ a t ) A~ TABLE 8 ( C o n t . ) LOCAL NUCLEATE BOILING HEAT TRANSFER RESULTS Run T ~ a t G ~g L~~ q l t x NO. h~~ ( T ~ - T ~ a t ) A~ TABLE 8 ( C o n t . ) LOCAL NUCLEATE BOILING HEAT TRANSFER RESULTS Run No. TABLE 9 CRITICAL HEAT FLUX RESULTS Run No. D a t e T i m e H g T ~ a t c s"c l o -5 A~, G X 1 12-29-67 1350 81.52 975 0.246 4.5 0.26 2 12-29-67 1800 115 1060 0.6 3.06 6.1 3 12-29-67 1850 112.6 1060 0.56 2.97 5.1 4 12-29-67 2005 158 1110 0.544 2.28 17.2 5 12-29-67 2130 1 4 1 1100 0.537 2.67 5.8 6 12-29-67 2230 176 1120 0.754 2.4 17.75 7 12-29-67 2255 178 1110 0.728 2.3 17.1 8 12-29-67 2315 163 1100 0.77 2.58 15.9 9 1-03-68 1145 191.4 1110 0.637 1.86 15.05 10 1-03-68 1300 185 1120 0.64 2.05 14.15 11 1-03-68 1500 186 1120 0.587 2.16 12.15 12 1-03-68 1900 193 1120 0.643 2.12 15.50 1 3 1-06-68 2300 192 1130 0.81 2.2 24.3 14 1-07-68 2110 193 1110 0.78 1.96 22.6 TABLE 1 0 LOCAL TRANSITION BOILING HEAT TRANSFER RESULTS Run No. G ~g *sat q;B X TB h~~ TABLE 10 (Cont.1 LOCAL TRANSITION BOILING HEAT TRANSFER RESULTS Run No. G ~g T ~ a t q;B X T B 'TB TABLE 11 LOCAL SUPERHEATED VAPOR RESULTS Run No. TABLE 11 (Cont. ) LOCAL SUPERHEATED VAPOR RESULTS Run No. G T ~ a t L Hg SH TABLE 11 (Cont .) LOCAL SUPERHEATED VAPOR RESULTS Run No. G ~g T ~ a t L h~~ TABLE 11 (Cont .) LOCAL SUPERHEATED VAPOR RESULTS Run No, G T ~ a t L ~g h~~ TABLE 1 2 LOCAL PRESSURE DROP DATA I N TWO-PHASE REGION Run G NO. ~g T ~ a t Sat T P ~ $H @M TABLE 1 2 (Cont.) LOCAL PRESSURE DROP DATA I N TWO-PHASE REGION Run ** G T ~ a t NO. ~ $ 4 S a t T P ~ T P ~ TP f +H +M 0 r . a r . m m . . . O C O 0 m w m w 9 . . m u m o m w ~ m m o w w m r l r . ON^ . . . 0 0 0 m m w " ? S 4 4 N m w w 4 m w . . . 0 0 0 0 m m o u e r . m TABLE 1 2 (Cont.) LOCAL PRESSURE DROP DATA I N TWO-PHASE REGION ** Run Fg T ~ a t No. S a t TPm T P f +H +M % 0, d w a 3 4 I- ln N N In a . 0 0 v) 0 d a 0 d 0 d ln 0 d w l n 9 "5 d rl ,-I ln m a3 * . 0 d ln m ? P 0 0 o m m + . . 0 0 w v) 0 N W . . 0 0 d 00 N N . . d N a3 v) d A d N APPENDIX 111 SNAP-8 SINGLE TUBE BOILER DATA REDUCTION RESULTS COMPUTER OUTPUT TABLE NOMENCLATIJRF! CATSB Reference Temperature Box "B" Cu-Const. T/C CATSC Reference Temperature Box "c" Cu-Const . T JC DTPPEX Pinch-point Temperature Di f fe rence a t B o i l e r Ex i t DTSH Degree of Superheat a t B o i l e r Ex i t GWR Axial Mercury Mass Veloc i ty i n t h e He l ix Region G.4WC Axial Mercury Mass Veloc i ty i n t h e Wire Co i l Region OPRl Mass Veloci ty of NaK Flow GSEC S u p e r f i c i a l Mass Veloc i ty of Mercury Flow NREES Superf i c i a1 Reynolds Number P3 B o i l e r Exi t Pressure P5 Condenser Pressure , Top Drum P6 Pump Exi t P re s su re P8 B o i l e r I n l e t P re s su re QB Heat Transfer red a f t e r B o i l i n g Incep t ion QNET Net Thermal Power t o B o i l e r Tube, QSc Heat Transfer red i n Subcooled Liquid Mercury QT T o t a l Thermal Power t o B o i l e r Q./ A Average Tes t Sect i o n Heat F lux T (1 ,2 , B o i l e r S h e l l T/C ( s t a t i o n number) -C i r cu l a r Locat ion 4 ,5 ,7 ,8 , l O , l l , 13? 14,16.17, 19,20,21, 22,231-A, B,C,D.E. TBS( 1,2--23) Average S h e l l Temperature TAMB Ambient Temperature TBEND Ex i t Bend Mercury Skin Temperature ( 1 , 2 ) T ~ J , B o i l e r Exi t Well Temperature l b see- f t 2 p s i p s i p s i psi TBIW (1,2) B o i l e r I n l e t W e l l Temperature TI3 I B o i l e r I n s e r t T/C ( S t a t i o n Number) (1 ,3 ,6 ,9 , TCLDW TCUDW TESAT TEXBHG TFMD TGH IN TINY TNAKI (1:2) TNAKE (1,2) TOUTW TPEW TPG 1 TPG2 (HI .LO) TPG3 TPG4 (HI ,LO) TPG7 (HI, LO) TPG8 TPHIW TS ING T SNAKX Condenser Lower D r u m Wall Temperature Condenser Upper D r u m Wall Temperature B o i l e r Ex i t S a t u r a t i o n Temperature Mercury Temperature a t B o i l e r Exi t NaK Flowmeter Duct Temperature NaK Flowmet e r Magnet Temperature Gas Bea t e r I n l e t Temperature B o i l e r I n l e t Wall Temperature NaK Inlet Well Temperature NaK Exi t Well Temperature Boi l e r Out l e t Wall Temperature Fvmp Ex i t Well Temperature Gas Heater I n l e t Pressure Transducer Temperature NaK Pres su re Gage Temperature B o i l e r Ex i t P re s su re Transducer Temperature Vapor Ventur i High and low Pres su re Tranducer Temp. Condenser Top Drum Pres su re Transducer Temp. O F Pump Exi t P re s su re Transducer Temperature " F Liquid Ventur i High and Low Pressure Transducer Temp. O F B o i l e r I n l e t P re s su re Transducer Temperature O F P r ehea t e r I n l e t Well Temperature O F Mercury I n l e t S a t u r a t i o n Temperature F S t a t i c NaK Expansion Tank Temperature OF Mercury Flow Rate (Liquid Ventur i ) l b / s ec WFAK NaK Flow Rate (NaK Flowmeter) Ib see X Quality - DATE 7 t M E DRW8R S N A ? Q @ SIWGLt*TUBE B O l b t R UPYA WEUUGTEON r l 9 ~ T 2 0 A ~ 3 0 8 T 2 0 C 721A * W M P W W H W H W P F W P e e B - - - - - 8 0 s - 8 IU G J N W fA K A C A fV W W W fA G 4 C d C d W 9 F c D 9 I U W w m a fa r a ~ C 3 S N A u l \a W *r B I* . B w w .1 \w m o e m w a w u e ~ - - a m ~ W A N O rn m m m m m m m r n r n m m m m m m s a ~ ~ ~ n s - 9 ) - 8 w c l , W C S d W 'L d ;+ 9 ip & e 9 & .Z x 2 a w * * a a w m & e ITI m in rr t m ' I 1, Gordon, R. and S lone , H . 0 . S N ~ I P - ~ Uevelopment S t a t u s . " AI,?,A S p e c i a l i s t s Conference on Rankice Space Power Systems, pp, 102-138, 1965, 2 . Thur, G . M. "SNAP-8 Power Conversion System Assessment. ' ' ~ ~ r s o c i e t y E n e r m Conversion Engineer ing C o ~ f e r e n c e , pp. 329-337, 1963. 3 , Brooks, R. D. SNAP-8 R e f r a c t o r v 3 3 i l e r Development Monthly P r o g r e s s Report . Genera1 E l e c t r i c L'o=?c-;.-, SPPS-FED, C o n t r a c t !;AS 3-10610, Augus t 29 6 7. Poppendiek, H. F. I n v e s ~ i g a ~ i o r . of t h e B o i l e r C o n d i t i o n i r . ~ and Heat T r a n s f e r C h a r a c t e r i s t i c s 2 z ?!erccrv-Tantalum System N o r t h l v T e c h n i c a l P r o g r e s s Report . Geosc ience , L 9 ; A e r o j e t C o n t r a c t Number 0P l l9224 , October 1967. 5. Bond, J. A. and Converse, G, I. : - a ~ o r i z a t i o n o f High T e x e r a t u r e 0 Potass ium in Forced Convect ion at S a t u r a t i o n Temperatures &.180(11 F t o 2 . 1 0 0 ~ ~ . NASA CR-843, .July 1°57. - 6. P e t e r s o n , J. R. High Performance Once-Throukh B o i l i n g o f S i n g l e Tubes a t S a t u r a t i o n T e z ~ e r a r u r e s of 1 5 1 0 0 ~ ~ to CR-842, August 1967. 7. Brooks, R. D. SNAP-8 Ref rac to r . - E o i l e r Development MonthlyL- P r o g r e ~ Report . General E l e c t r i c Cor_?zry, SPPS-MSD; C o n t r a c t NAS 3-10610; October , November, and Decsmber 1967. 8. Chang, Y. P. "Wave Theory o f Fez= T r a n s f e r i n Film E o i l i - g , " J o u r n a l o f Heat T r a n s f e r , Vol. 81, feSruzr:- 1959. -- 9 , Zuber, N. Hydrodynamic ,Aspects cf B o i l i n g Heat T r a n s f e r . ? h . D , T h e s i s , U n i v e r s i t y of C a l i f o r n i a , Los Azgeles , June 1959. - - 10. Berenson, P, J. "Film B o i l i n g zez: T r a n s f e r from a H o r i z o n t a l S u r f a c e , " J o u r n a l of F e a t T r a n s f e r , Vol. 83 , No. 3, 1961. 11, S e l l e r s , A. J. ; Thur, G. ?I. ; an? Xong, M. K. Recent D e s ~ e l o c m e z ji Heat T r a n s f e r 4 Developre r+ o f t h e Mercury B o i l e r for 2 SNAP-8 - -- S y s t e m Aeroje t -General Co~i)3ra:ion, SNAP-8 D i v i s i o n , 1°65, 12.. " P r o p e r t i e s of F lu ids . " Design :.!anual H-100. Aeroje t -Gezi i ra l Corpora t ion , January 1967, 13. Converse, G . L. and H s i a , E. S. ; I - e m a l e H y d r a u l i c A n a l v s i s for Advanced SNAP-8 B o i l e r s , General E l e c t r i c Company, SPPS 9TC-14, January 1969. 14. Lyon, R. N. "Liquid Meta l Peat T r a n s f e r C o e f f i c i e n t s . " i r z n s , , ATChE, 47:75-79, 1965. REFERENCES (Cont'd) 15. Lubarsky, B. and Kaufmann, S. J. Review of Exper imenta l I n v e s t i g a t i o n of Liquid-Metal Heat T r a n s f e r . NASA Report No. 1270, 1956. - 16. Atomics I n t e r n a t i o n a l . SNAP-8 Systems Improvement Program P r o g r e s s Report .July-September 1966, Vol. I. NAA-SR-12256, January 1967. 17. K u t a t e l a d z e , S. S. L i q u i d Metal Hea t T r a n s f e r Media. New York: C o n s u l t a n t s Bureau, I n c , , 1959. 18. C i c h e l l i , M. T. and B o n i l l a , C. F. Trans. AIChE, 41:755-787, 1945. 19. Zuber, N. " S t a b i l i t y o f B o i l i n g Heat ~ r a n s f e r . " Trans . ASPIE, Vol. 80-C, A p r i l 1958. 20. Merte, J., Jr. and Cla rk , J. A, S tudy o f Poo l oiling i n an Acce le ra t inq , System. ASME Paper 60-HT-22, 1960. 21. Longo, J. A l k a l i Meta l B o i l i n g and Condensing I n v e s t i g a t i o n Q u a r t e r l y P r o g r e s s Report No. 6. Genera l E l e c t r i c Company, SPPS, C o n t r a c t NAS 3-2528, A p r i l 1964. 22. Brooks, R. D. A l k a l i M e t a l B o i l i n g & Condensing I n v e s t i g a t i o n , Vol. I. Genera l E l e c t r i c Company, SPPS, C o n t r a c t NAS 5-681, J u n e 1964. 23. Levy, S. "Steam S l i p - T h e o r e t i c a l P r e d i c t i o n from Momentum ~ o d e l , " J o u r n a l of Hea t T r a n s f e r , 82:113, May 1960. 24. K u t a t e l a d z e , S. S, Fundamentals o f Hea t T r a n s f e r . New York: Academic P r e s s , I n c . , 1963, 25. C o s t e l l o , C. P. and Adams, J. M. "Burnout Heat F luxes i n Poo l B o i l i n g a t High Acce le ra t ions . " I n t e r n a t i o n a l Heat T r a n s f e r Conference, Paper No. 30, August 1961. 26. Veiskim, C. M. and S i e g e l , R. "An Exper imental Study o f B o i l i n g i n Reduced and Zero-Gravity ~ i e l d s . " Trans. ASME, 83-C:243-253, August 1961. 27. Converse, G , L. & Heat ing Experiments. Genera l E l e c t r i c Company, SPPS HTC-12, A p r i l 1969. 28. Sams, E. W. "Heat T r a n s f e r and P r e s s u r e Drop C h a r a c t e r i s t i c s of Wire-Coil Type Turbulence Promoters." R e a c t o r Hea t T r a n s f e r Conference. U.S. AEC, November 1956. 29. Converse, G. L. and Tedesco, D. P. A i r Water Experiment - T r a n s p a r e n t P l a s t i c Tube. Genera l E l e c t r i c Company, SPPS HTC-10, March 1968. 30. Hsu, C. J. "Heat T r a n s f e r i n a Round Tube w i t h S i n u s o i d a l Wall Heat Flux D i s t r i b u t i o n . " AAIhE J o u r n a l , 11:690-695, 1965. REFERENCES (Cont ' d) 31. S t e i n , R. P. "Improved Design P r e d i c t i o n s Through Liquid-Metal Heat T r a n s f e r Research." Power Reac to r Technology, 9:174-179, 1966, 32. Nunge, R. J. and G i l l , W. N. "An A n a l y t i c a l Study o f Laminar Counter- f low Double-Pipe Heat Exchangers." AIChE J o u r n a l , 12:279-289, 1966. 33. Dwyer, 0. E. "Heat T r a n s f e r t o F l u i d s Flowing Through P i p e s , Annuli and P a r a l l e l P l a t e s " USAEC Report No. BNL-6692, 1963. 34. M a r t i n e l l i , R. C. and Nelson, D. B. " P r e d i c t i o n o f P r e s s u r e Drop During Forced-Ci rcu la t ion B o i l i n g o f water." Trans . ASME, 7 0 ~ 6 9 5 , 1948. 35. Fauske, H, K. C o n t r i b u t i o n t o t h e Theory of Two-Phase, C r i t i c a l Flow. ANL-6633, October 1962. 36. M a r t i n e l l i , R. C.; B o e l t e r , L. M. K.; Tay lor , T. H. M.; Thompson, E, G.; and Morr in , E. H. '"sothermal P r e s s u r e Drop f o r Two-Phase, 73.70-Component Flow i n a H o r i z o n t a l Pipe." Trans . ASME, 66:139, 1944. 37. Lockhar t , R, W. and M a r t i n e l l i , R, C. ' 'Proposed C o r r e l a t i o n of Data for I s o t h e r m a l Two-Phase Two-Component Flow i n ~ i ~ e s . " Chem. Prog. , 4 5 ~ 3 9 , 1949, 1 ~ 1 S T R I J ~ ~ ' ' ~ I i ) ~ ~ L?!;T '17 1P 1 (XI , ANT) FZ!;,\L RIII'URTS :!A!; 7-10619 ; ; ; i t i onn l :?cror.:lutLcs and Sp'icc. Aclminiscration ~ ~ : a s h i n ~ : t o n , D. C. 20546 A t t e : ~ t i o n : P .X. Y i ' l l e r (RN?') .Tr?mes .J. I,?nch (K?;P) Gc6-rge C. Deutsch (m) W. ?,',ochen (RSP) J. Cangle r N a t i o n a l A e r o n z u t i c s and Space Adminis:r=tion Lewis Zesearc5 Cente r 21000 Brookpark Zoad Cleveland, Ohio 44135 A t t e n t i o n : L i b r a r i a n G.'!. A u l t , ?!S 3-13 P.L. 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Livermore, C a l i f o r n i a 94550 A t t e n t i o n : L i b r a r i a n Lockheed M i s s i l e s and Space D i v i s i o n Lockheed A i r c r a f t C o r p o r a t i o n Sunnyvale, C a l i f o r n i a 90221 A t t e n t i o n : L i b r a r i a n Los Alamos S c i e n t i f i c Labora to ry U n i v e r s i t y of C a l i f o r n i a Los Alamos, New Mexico A t t e n t i o n : L i b r a r i a n Ela ter ia l s Research & Development I'ianla'bs , Incorporated 2 1 E r i e S t r e e t Cambridge, Massachusetts 02139 Flarquardt A i r c r a f t Co, P. 0 . Box 2013 Van Nuys, C a l i f o r n i a At ten t ion: L ib ra r i an The Mart in Company ;Juclear Div is ion F. 0 , Box 5042 Xaltirnore, Maryland 21203 Atten t ion: L ib ra r i an Mart in 2 l a r i e t t a Corporation :Ietal:s Technology Laboratory 'IrYheel:ing, I l l i n o i s 60090 f la ter : ia ls Research Corporation Orangeburg, New York 10962 A t t en t ion : L ib ra r i an PIcDonnell A i r c r a f t S t , L o u i s , Missouri 03166 At ten t ion: L ib ra r i an t1SA Research Corporation Calle:ry, Pennsylvania 16024 A t t en t ion : L ib ra r i an National Research Corporation 70 Mernorial Drive Cambridge, Massachuset ts 02142 A t t en t ion : L i b r a r i a n North American Rockwell Los Angeles Div is ion Los A ~ ~ g e l e s , C a l i f o r n i a 90009 Attenition: L ib ra r i an North American Rockwell Corporation Atomics I n t e r n a t i o n a l Div is ion 8900 DeSota Avenue Canoga Park, C a l i f o r n i a 91304 Attent ion: T.A. Moss P.B, Fer ry L i b r a r i a n Ph i l co Corporation Aeronrstronics Newpo:rt Beach, C a l i f o r n i a 92665 A t t en t ion : L i b r a r i a n Republic Avi Famingda le , At ten t ion : . a t i o n Corporat ion Long I s l and , New York 11735 L i b r a r i a n Kocketdyne Canoga Park, C a l i f o r n i a 91303 At t en t ion : L i b r a r i a n Sandia Corporat ion P. 0. Box 5800 Albuquerque, New Mexico 87116 At ten t ion : L i b r a r i a n Don Johnson So la r 2200 P a c i f i c Highway San Diego, C a l i f o r n i a 92112 At ten t ion : L i b r a r i a n Southwest Research I n s t i t u t e 8500 Culebra Road San Antonio, Texas 78228 At ten t ion : L i b r a r i a n Super ior Tube Company Norristown, Pennsylvania 19404 At ten t ion : L i b r a r i a n TRW, Inc. 23555 Eucl id Avenue Cleveland, Ohio 44117 At ten t ion : L i b r a r i a n K. S h e f f l e r Union Carbide Corporat ion Ma te r i a l s Systems Div is ion 1020 West Park Kokomo, Ind iana 46901 At ten t ion : L i b r a r i a n Technology Department Union Carbide Metals Niagara F a l l s , New York 14300 At ten t ion : L i b r a r i a n Union Carbide Corporation Parma Research Center P. 0. Box 6115 Cleveland, Ohio 44101 At ten t ion : Technical Info. Serv ices United A i r c r a f t Corporation P r a t t & Whitney A i r c r a f t Div is ion 400 Main S t r e e t Eas t Hartford, Connecticut 06108 At ten t ion: L ib ra r i an TJah Chang Corporation Albany, Oregon 97321 A t t en t ion : L i b r a r i a n \,?estinghouse E l e c t r i c Corporation Astronuclear Laboratory P. 0. Box 10864 P i t t sbu rgh , Pennsylvania 15236 At ten t ion: R.W. Ruckman D.R, Stoner L ib ra r i an Westinghouse E l e c t r i c Corporation Research & Development Center P i t t sbu rgh , Pennsylvania 15235 At ten t ion: L i b r a r i a n R.T, Begley Westinghouse E l e c t r i c Corporation Mate r i a l s Manufacturing Div is ion RD i f2, Box 25 B l a i r s v i l l e , Pennsylvania At ten t ion: L i b r a r i a n

Aerojet-General Nucleonics mercury corrosion loop pump

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Aerojet-General Nucleonics mercury corrosion loop pump

Abstract

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