UBC Theses and Dissertations

UBC Theses Logo

UBC Theses and Dissertations

Methods for the characterization of deposition and transport of magnetite particles in supercritical… Karakama, Keigo 2011

Your browser doesn't seem to have a PDF viewer, please download the PDF to view this item.

Item Metadata

Download

Media
24-ubc_2012_spring_karakama_keigo.pdf [ 12.46MB ]
Metadata
JSON: 24-1.0072444.json
JSON-LD: 24-1.0072444-ld.json
RDF/XML (Pretty): 24-1.0072444-rdf.xml
RDF/JSON: 24-1.0072444-rdf.json
Turtle: 24-1.0072444-turtle.txt
N-Triples: 24-1.0072444-rdf-ntriples.txt
Original Record: 24-1.0072444-source.json
Full Text
24-1.0072444-fulltext.txt
Citation
24-1.0072444.ris

Full Text

Methods for the Characterization of Deposition and Transport of Magnetite Particles in Supercritical Water  by Keigo Karakama  B.A.Sc., The University of British Columbia, 2009  A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF  MASTER OF APPLIED SCIENCE in THE FACULTY OF GRADUATE STUDIES (Mechanical Engineering)  THE UNIVERSITY OF BRITISH COLUMBIA (Vancouver)  December 2011  © Keigo Karakama, 2011  Abstract Generation IV CANDU Supercritical Water Reactor (SCWR) is being developed to use a light water coolant at high temperature and pressure beyond the critical point of water (374⁰C and 22.1 MPa). The dramatic decrease in the solubility of magnetite in supercritical water suggests that the precipitation of magnetite particles will occur in the reactor core which can deposit on the fuel cladding or be transported to the steam turbine.  A once-through flow system was modified to develop experimental techniques for studying the deposition and transport of magnetite particles in supercritical water onto stainless steel 316L. Experiments were run with temperatures ranging from 200°C to 400°C and a pressure of 23.7 MPa. A hydrothermal method for synthesizing magnetite particles was adapted for producing simulated corrosion products in which a typical run had an iron concentration of 0.005 mol/L and lasted for 40 minutes.  An online monitoring technique using thermal resistance to infer deposit loadings showed deposition and removal cycles of the corrosion product on the tube wall. Scanning electron microscope images of particles on the tube inner wall and those collected by the high temperature, high pressure filters revealed magnetite particles which were several hundred nanometers to several microns in diameter depending on the precursor and condition of the system. Ultrasound and acid wash cleaning methods were used to remove deposits from the test section for determining deposit thickness and adhesive strength. The strength of deposit adhesion was observed to increase along the tube, particularly under supercritical conditions suggesting precipitation of dissolved species may enhance the strength of the deposit. By  ii  comparing the results, a comprehensive approach was developed to study magnetite fouling in supercritical water conditions.  Finally, comparison between a simulation model based on mass transport equations and experimental deposition suggests that mass transport alone can overestimate the deposition thickness when surface attachment and removal are significant as they were for many experiments in this study. The simulation predicted as an upper limit scenario that fouling in a CANDU SCWR could increase the fuel cladding temperature at certain locations by up to 23.9⁰C after one year of operation.  iii  Preface Chapter 5 is based on experimental work conducted at the mechanical engineering laboratories by the author, Keigo Karakama. Thermocouple temperature measurements and its analysis were done by the author. All Atomic absorption spectroscopy analysis, scanning electron microscopy images and energy dispersive x-ray analysis was completed by the author at the Materials Engineering Laboratories. X-ray diffraction data and analysis in Chapter 5 was produced by Anita Lam in the Chemistry Department.  iv  Table of Contents Abstract ..................................................................................................................................... ii Preface...................................................................................................................................... iv Table of Contents ...................................................................................................................... v List of Tables ......................................................................................................................... viii List of Figures .......................................................................................................................... ix List of Symbols and Abbreviations......................................................................................... xv Acknowledgements ............................................................................................................... xvii 1  2  Introduction ....................................................................................................................... 1 1.1  Supercritical Water ..................................................................................................... 1  1.2  Systems using Supercritical Water Medium .............................................................. 3  1.3  Scope and Objectives ................................................................................................. 3  Literature Review.............................................................................................................. 6 2.1  3  Applications of Supercritical Water ........................................................................... 6  2.1.1  Supercritical Fossil-fueled Power Plants ............................................................ 6  2.1.2  Supercritical Water Oxidation ............................................................................ 8  2.1.3  Nuclear Energy and the Generation IV Nuclear Reactors .................................. 8  2.1.4  CANDU Supercritical Water Reactors ............................................................. 10  2.2  Materials for SCWR ................................................................................................. 11  2.3  Corrosion and Fouling in Subcritical and Supercritical Water ................................ 13  2.3.1  Properties of Magnetite ..................................................................................... 13  2.3.2  Magnetite and its Solubility in High Temperature and Pressure Water ........... 15  2.3.3  Corrosion........................................................................................................... 17  2.3.4  Fouling .............................................................................................................. 18  2.3.5  Corrosion Product Removal and Transport ...................................................... 22  Heat Transfer and Deposition Modeling......................................................................... 25 3.1  Overview .................................................................................................................. 25  3.2  Heat Transfer Calculations of Water........................................................................ 25  3.2.1  Enthalpy of Fluid in the Test Section ............................................................... 25  3.2.2  Heat Transfer Coefficient in Supercritical Water ............................................. 26  v  4  3.2.3  Temperature of the Oxide Layer ....................................................................... 28  3.2.4  Temperature of the Outside Wall Temperature ................................................ 28  3.2.5  Buoyancy Calculations ..................................................................................... 29  3.3  Deposition Calculations ........................................................................................... 31  3.4  CANDU Supercritical Water Reactor Simulation ................................................... 38  Experiments .................................................................................................................... 41 4.1  Supercritical Water Once-through Flow Apparatus ................................................. 41  4.2  Flow and Temperature Control ................................................................................ 44  4.2.1  Safety ................................................................................................................ 44  4.2.2  Pressure Measurements ..................................................................................... 45  4.2.3  Flow Rate Measurements .................................................................................. 45  4.2.4  Fluid Conductivity Measurements .................................................................... 46  4.2.5  Dissolved Oxygen Concentration Measurements ............................................. 46  4.3  Particle Synthesis for Modeling Power Plant Fouling ............................................. 46  4.4  Experimental Procedure ........................................................................................... 48  4.4.1  Experiments with Precursor .............................................................................. 48  4.4.2  Experiments with Magnetite Particles .............................................................. 50  4.4.3  Post Experiment Cleaning................................................................................. 51  4.5  5  Analytical Methods .................................................................................................. 51  4.5.1  Online Thermal Resistance Monitoring ............................................................ 51  4.5.2  SEM Imaging for Filters and Tube Deposits .................................................... 59  4.5.3  Particulate Filtering System .............................................................................. 60  4.5.4  Deposit Thickness and Strength of Oxide Adhesion ........................................ 67  Results and Discussion of Experiments .......................................................................... 69 5.1  Reference Condition ................................................................................................. 70  5.1.1  Temperature Analysis ....................................................................................... 70  5.1.2  Filter Analysis ................................................................................................... 72  5.1.3  Deposit Analysis ............................................................................................... 73  5.2  Effect of Heat Flux ................................................................................................... 80  5.2.1  Temperature Analysis ....................................................................................... 80  5.2.2  Filter Analysis ................................................................................................... 80  vi  5.2.3 5.3  6  Deposit Analysis ............................................................................................... 84  High pH Conditions ................................................................................................. 87  5.3.1  Temperature Analysis ....................................................................................... 89  5.3.2  Filter Analysis ................................................................................................... 91  5.3.3  Deposit Analysis ............................................................................................... 93  5.4  Low Concentration and Blank Experiment .............................................................. 95  5.5  Overall Summary and Mass Balance ....................................................................... 95  Results and Discussion of Simulations ......................................................................... 100 6.1  Comparison between Simulation and Experimental .............................................. 100  6.2  Simulation for CANDU SCWR ............................................................................. 103  6.3  Limitations of the Simulation................................................................................. 107  7  Conclusion .................................................................................................................... 109  8  Recommendations ......................................................................................................... 112  References ............................................................................................................................. 113 Appendix A: MATLAB Code .............................................................................................. 120 Appendix B: Filter Parts Drawings ....................................................................................... 131 Appendix C: Results of Experiments.................................................................................... 133 C. 1.  Reference Condition, Ferrous Chloride .............................................................. 133  C. 2.  Reference Condition, Ferrous Sulfate ................................................................ 138  C. 3.  Low Heat Flux, Ferrous Chloride ....................................................................... 143  C. 4.  Low Heat Flux, Ferrous Sulfate ......................................................................... 148  C. 5.  No Heat Flux, Ferrous Sulfate ............................................................................ 153  C. 6.  High pH, Ferrous Sulfate .................................................................................... 158  C. 7.  High pH & Subcritical, Ferrous Sulfate ............................................................. 163  C. 8.  Low Concentration, Ferrous Chloride ................................................................ 166  C. 9.  Blank Run ........................................................................................................... 169  Appendix D: Filter Flow Rates ............................................................................................. 172 Appendix E: Atomic Absorption Spectroscopy.................................................................... 173  vii  List of Tables Table 2-1: Generation IV systems and expected deployment year........................................... 9 Table 2-2: Stainless steel AISI 316L composition (American Society for Metals, 1985) ..... 12 Table 3-1: SCWR and fuel assembly design parameters ........................................................ 39 Table 4-1: Input power comparison ........................................................................................ 55 Table 4-2: Glass and silver membrane filter comparison ....................................................... 63 Table 5-1: Summary of experiments....................................................................................... 69 Table 5-2: Results summary table - Part I .............................................................................. 96 Table 5-3: Results summary table - Part II ............................................................................. 97 Table 5-4: Mass balance of experiments #1 - 8 ...................................................................... 99 Table C-1: Summary table of experiment #1 ........................................................................ 133 Table C-2: Summary table of experiment #2 ........................................................................ 138 Table C-3: Summary table of experiment #3 ........................................................................ 143 Table C-4: Summary table of experiment #4 ........................................................................ 148 Table C-5: Summary table of experiment #5 ........................................................................ 153 Table C-6: Summary table of experiment #6 ........................................................................ 158 Table C-7: Summary table of experiment #7 ........................................................................ 163 Table C-8: Summary table of experiment #8 ........................................................................ 166 Table C-9: Summary table of blank run ............................................................................... 169 Table D-1: Flow rate through filter ....................................................................................... 172  viii  List of Figures Figure 1-1: Pressure - temperature diagram for water, grey region indicating supercritical conditions............................................................................................................. 1 Figure 1-2: Density of water vs. temperature at various pressures, showing the effects of pressure on the pseudocritical temperature ......................................................... 2 Figure 2-1: Simplified drawing of a direct cycle coolant for Generation IV CANDU SCWR ........................................................................................................................... 11 Figure 2-2: Pourbaix diagram (a) Fe – H2O at 25ºC and 1 atm, (b) Fe – H2O at 374°C and 220 atm (c) Fe/Cr – H2O at 25°C and 1 atm, (d) Fe/Cr – H2O at 374ºC and 220 atm. Based on SUPCRT92 (Propp et al., 1996) ............................................... 14 Figure 2-3: Magnetite solubility adapted from Cook and Fatoux (2009) for neutral pH water at 25 MPa, Guzonas et al. (2009) for pH = 9.3 at 30 MPa, Burrill (2000) for neutral pH at 25 MPa, and Sweeton and Baes Jr. (1970) experimental results. 16 Figure 3-1: Calculation of Grq/Grth for determination of the effects of buoyancy along the test section using Bazargan et al. (2005) ........................................................... 31 Figure 3-2: Effect of particle diameter and fluid temperature on deposition velocity for experimental apparatus ...................................................................................... 38 Figure 3-3: Cross-sectional view of CANFLEX 43 rod fuel bundle and drawing of equivalent tube for simulation calculations (not to scale) ................................. 40 Figure 4-1: SCW system used in all experiments located in the Mechanical Engineering laboratories. ....................................................................................................... 41 Figure 4-2: Schematic of supercritical once-through flow apparatus................................... 42 Figure 4-3: Mounting block and test section inlet ................................................................ 54 Figure 4-4: Electrical wiring, heater controls and location of surface welded thermocouples and bulk fluid thermocouples ............................................................................ 57 Figure 4-5: Saturation temperature of 231.51ºC at 417psi, calibration for Experiment #1 . 58 Figure 4-6: Temperature of test section, calculated vs. calibrated thermocouples............... 58 Figure 4-7: Test section sample for SEM analysis ............................................................... 60 Figure 4-8: SS316L LTHP filter installed after the heat exchanger ..................................... 61 Figure 4-9: HTHP filter with silver membrane and SS304 back support ............................ 62 Figure 4-10: Glass fiber membrane and Swin-Lok LTLP filter holder .................................. 63 Figure 4-11: SEM photograph of clean silver membrane, 0.2µm pore size, using a Hitachi S3000N SEM ..................................................................................................... 65 Figure 4-12: EDX of silver membrane to determine elemental composition of background 65 Figure 4-13: SEM photograph of clean glass fiber filter, 0.7µm pore size, a Hitachi S2300 SEM ................................................................................................................... 66 Figure 4-14: EDX of glass fiber filter to determine elemental composition of background .. 66 Figure 5-1: Temperature measurements of test section vs. time for reference condition with ferrous chloride precursor (Experiment #1) ...................................................... 71  ix  Figure 5-2: SEM photograph of magnetite deposit on test section tube surface, 0.15 m location. Reference condition with ferrous chloride precursor (Experiment #1) ........................................................................................................................... 73 Figure 5-3: SEM photograph of magnetite deposit on test section tube surface, 0.60 m location. Reference condition with ferrous chloride precursor (Experiment #1) ........................................................................................................................... 74 Figure 5-4: SEM photograph of magnetite deposit on test section tube surface, 1.20 m location. Reference condition with ferrous chloride precursor (Experiment #1) ........................................................................................................................... 74 Figure 5-5: SEM photograph of magnetite deposit on test section tube surface, 1.65 m location. Reference condition with ferrous chloride precursor (Experiment #1) ........................................................................................................................... 75 Figure 5-6: XRD of tube deposits of test section, reference condition ferrous chloride precursor (Experment #1) .................................................................................. 76 Figure 5-7: XRD of tube deposit of test section, reference condition ferrous sulfate precursor (Experiment #2) ................................................................................. 76 Figure 5-8: Deposit thickness on test section for reference condition with ferrous chloride precursor (Experiment #1) ................................................................................. 79 Figure 5-9: Deposit thickness on test section for reference condition with ferrous sulfate precursor (Experiment #2) ................................................................................. 79 Figure 5-10: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for low heat flux with ferrous chloride precursor (Experiment #3) ........................ 81 Figure 5-11: EDX of HTHP filter for low heat flux with ferrous chloride precursor ............ 81 Figure 5-12: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for low heat flux with ferrous sulfate precursor (Experiment #4)........................... 82 Figure 5-13: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for no heat flux with ferrous chloride (Experiment #5) .......................................... 82 Figure 5-14: XRD of deposit on HTHP filter for low heat flux with ferrous chloride precursor (Experiment #3) ................................................................................. 83 Figure 5-15: Deposit thickness on test section, low heat flux with ferrous sulfate precursor (Experiment #4) ................................................................................................. 84 Figure 5-16: Calibration of the conductivity meter for determining NaOH concentration in primary tank ....................................................................................................... 88 Figure 5-17: Predicted pH of the system using data from Arden (1950).............................. 89 Figure 5-18: Temperature measurements of test section vs. time for high pH with ferrous sulfate precursor (Experiment #6). Power turned off briefly at t = 48 minutes. 90 Figure 5-19: SEM photograph of particles collected on HTHP filter, 10.0k magnification, for high pH, ferrous sulfate (Experiment #6) .......................................................... 91 Figure 5-20: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH, ferrous sulfate (Experiment #6) .......................................................... 92  x  Figure 5-21: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for high pH, subcritical, ferrous sulfate (Experiment #7) ....................................... 92 Figure 5-22: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH, subcritical, ferrous sulfate (Experiment #7) ....................................... 93 Figure 5-23: Deposit thickness on test section (Experiment #6) ............................................ 94 Figure 6-1: Comparison between simulation and experimental for reference conditions (a) Experiment #1, ferrous chloride with average particle size = 4 µm, magnetite concentration = 200 mg/L (b) Experiment #2, ferrous sulfate with average particle size = 1 µm, magnetite concentration = 299 mg/L. ............................ 100 Figure 6-2: Comparison between simulation and experimental for ferrous sulfate experiments with (a) Experiment #4, low heat flux with average particle size = 1 µm, magnetite concentration = 108 mg/L and (b) Experiment #5, no heat flux with average particle size = 2 µm, magnetite concentration = 89 mg/L. ........ 101 Figure 6-3: Comparison between simulation and experimental for Experiment #6, high pH, ferrous sulfate. Simulation was run with an average particle size = 0.15µm, magnetite concentration = 321 mg/L. .............................................................. 102 Figure 6-4: Bulk fluid and fuel cladding-surface temperature without particle deposition 104 Figure 6-5: Bulk fluid and fuel cladding-surface temperature after 1 year of operation with magnetite particles of 150 nm diameter .......................................................... 104 Figure 6-6: Magnetite deposition thickness on fuel cladding after 1 year of operation for 150 nm and 1 µm average particle diameters .................................................. 105 Figure C-1: Temperature measurements of test section vs. time for reference condition with ferrous chloride precursor. (Experiment #1) ................................................... 134 Figure C-2: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for reference condition with ferrous chloride precursor. Low flow rate in secondary line resulting in very few particles collected. (Experiment #1) ....................... 134 Figure C-3: SEM photograph of particles collected on LTHP filter, 1.8k magnification, for reference condition with ferrous chloride precursor. (Experiment #1) ........... 135 Figure C-4: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for reference condition with ferrous chloride precursor. Low flow rate in secondary line resulting in very few particles collected. (Experiment #1) ....................... 135 Figure C-5: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #1) .................. 136 Figure C-6: Deposit thickness on test section for reference condition with ferrous chloride precursor. (Experiment #1) .............................................................................. 137 Figure C-7: Comparison between simulation and experimental for reference condition ferrous chloride. Simulation was run with an average particle size = 4 µm, magnetite concentration = 200 mg/L. .............................................................. 137  xi  Figure C-8: Temperature measurements of test section vs. time, temperature fluctuations were found in both the inlet and outlet bulk temperatures which may be attributed to the inconsistent flow. (Experiment #2) ....................................... 139 Figure C-9: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for reference condition with ferrous sulfate. (Experiment #2) .............................. 139 Figure C-10: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for reference condition with ferrous sulfate. (Experiment #2) .............................. 140 Figure C-11: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for reference condition with ferrous sulfate. (Experiment #2) .............................. 140 Figure C-12: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #2) .................. 141 Figure C-13: Deposit thickness on test section for reference condition with ferrous sulfate precursor. (Experiment #2) .............................................................................. 142 Figure C-14: Comparison between simulation and experimental for reference condition ferrous sulfate. Simulation was run with an average particle size = 1 µm, magnetite concentration = 190 mg/L. .............................................................. 142 Figure C-15: Temperature measurements of test section vs. time for low heat flux ferrous chloride precursor. (Experiment #3) ................................................................ 144 Figure C-16: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for low heat flux with ferrous chloride precursor. (Experiment #3) ..................... 144 Figure C-17: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for low heat flux with ferrous chloride precursor. (Experiment #3) ..................... 145 Figure C-18: SEM photograph of particles collected on LTLP filter, - magnification, for low heat flux with ferrous chloride precursor. (Experiment #3) ............................ 145 Figure C-19: SEM photograph of magnetite deposit on test section tube surface, (a) 0.10 m, (b) 0.30 m, and (c) 0.90 m location. (Experiment #3) ..................................... 146 Figure C-20: Deposit thickness on test section for low heat flux ferrous chloride precursor. (Experiment #3) ............................................................................................... 147 Figure C-21: Temperature measurements of test section vs. time for low heat flux with ferrous sulfate precursor. (Experiment #4) ...................................................... 149 Figure C-22: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for low heat flux ferrous sulfate precursor (Experiment #4) ................................. 149 Figure C-23: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for low heat flux ferrous sulfate precursor. (Experiment #4) ................................ 150 Figure C-24: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for low heat flux ferrous sulfate precursor. (Experiment #4) ................................ 150 Figure C-25: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #4) .................. 151 Figure C-26: Deposit thickness on test section for low heat flux ferrous sulfate precursor. (Experiment #4) ............................................................................................... 152  xii  Figure C-27: Comparison between simulation and experimental for low heat flux, ferrous sulfate. Simulation was run with an average particle size = 1 µm, magnetite concentration = 108 mg/L................................................................................ 152 Figure C-28: Temperature measurements of test section vs. time for no heat flux with ferrous sulfate precursor. (Experiment #5) .................................................................. 154 Figure C-29: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for no heat flux with ferrous sulfate precursor. (Experiment #5).......................... 154 Figure C-30: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for no heat flux with ferrous sulfate precursor. (Experiment #5).......................... 155 Figure C-31: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for no heat flux with ferrous sulfate precursor. (Experiment #5).......................... 155 Figure C-32: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #5) .................. 156 Figure C-33: Deposit thickness on test section for no heat flux with ferrous sulfate precursor. (Experiment #5) ............................................................................................... 157 Figure C-34: Comparison between simulation and experimental for no heat flux with ferrous sulfate precursor. Simulation was run with an average particle size = 2 µm, magnetite concentration = 89 mg/L. ................................................................ 157 Figure C-35: Temperature measurements of test section vs. time for high pH with ferrous sulfate precursor. (Experiment #6) .................................................................. 159 Figure C-36: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for high pH with ferrous sulfate precursor. (Experiment #6) ................................ 159 Figure C-37: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH with ferrous sulfate precursor. (Experiment #6) ................................ 160 Figure C-38: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for high pH with ferrous sulfate precursor. (Experiment #6) ................................ 160 Figure C-39: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #6) .................. 161 Figure C-40:Deposit thickness on test section for high pH with ferrous sulfate precursor. (Experiment #6) ............................................................................................... 162 Figure C-41: Comparison between simulation and experimental for high pH, ferrous sulfate. Simulation was run with an average particle size 0.15 µm, magnetite concentration = 321 mg/L................................................................................ 162 Figure C-42: Temperature measurements of test section vs. time for high pH & subcritical with ferrous sulfate precursor. (Experiment #7) .............................................. 164 Figure C-43: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for high pH & subcritical with ferrous sulfate precursor. (Experiment #7) .......... 164 Figure C-44: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH & subcritical with ferrous sulfate precursor. (Experiment #7) .......... 165  xiii  Figure C-45: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for high pH & subcritical with ferrous sulfate precursor. (Experiment #7) .......... 165 Figure C-46: Temperature measurements of test section vs. time for low concentration with ferrous chloride precursor. (Experiment #8) ................................................... 167 Figure C-47: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for low concentration with ferrous chloride precursor. (Experiment #8).............. 167 Figure C-48: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for low concentration with ferrous chloride precursor. (Experiment #8).............. 168 Figure C-49: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for low concentration with ferrous chloride precursor. (Experiment #8).............. 168 Figure C-50: Temperature measurements of test section vs. time for high pH & subcritical with ferrous sulfate precursor. (Experiment #9) .............................................. 170 Figure C-51: SEM photograph of particles collected on HTHP filter, 2.0k magnification, blank run. (Experiment #9) .............................................................................. 170 Figure C-52: SEM photograph of particles collected on LTHP filter, 5.0k magnification, blank run. (Experiment #9) .............................................................................. 171 Figure C-53: SEM photograph of particles collected on LTLP filter, 2.0k magnification, blank run. (Experiment #9) .............................................................................. 171 Figure E-1: Calibration curve for iron concentration using AAS, nitric acid matrix ......... 173 Figure E-2: Calibration curve for iron concentration using AAS, nitric acid and copper sulfate matrix ................................................................................................... 174  xiv  List of Symbols and Abbreviations AAS  Atomic Absorption Spectroscopy  AECL  Atomic Energy of Canada Limited  CANDU  Canadian Deuterium Uranium  Cb  Concentration of magnetite in bulk fluid  cp  Heat capacity [kJ/kg·K]  Csat  Saturation concentration of magnetite  dp  Particle diameter [m]  f  Friction factor  Gr  Grashoff number  H1  Enthalpy [kJ/kg]  H2  Enthalpy [kJ/kg]  HKF  Helgeson-Kirkham-Flowers  HTHP  High temperature high pressure  Ka  Surface attachment deposition velocity [m/s]  Kd  Effective deposition velocity [m/s]  Kt  Mass transport deposition velocity [m/s]  LTHP  Low temperature high pressure  LTLP  Low temperature low pressure  Nu  Nusselt number  ̇  Mass flow rate [kg/s]  Pr  Prandtl number  q  Heat input per meter [kW/m]  Re  Reynolds number  Sc  Schmidt number xv  SCWO  Supercritical Water Oxidation  SCWR  Supercritical Water Reactor  SEM  Scanning Electron Microscope  SPE  Solid Particle Erosion  Tb  Bulk fluid temperature [ºC]  Te  Temperature of outer wall [ºC]  tp +  Dimensionless particle relaxation time  Tw  Wall temperature [ºC]  U*  Frictional velocity  µ  Dynamic viscosity [Pa·s]  XRD  X-ray Diffraction  Δx  Interval length [m]  ε  Tube roughness  ρf  Density of fluid [kg/m3]  ρp  Density of particle [kg/m3]  τw  Wall shear stress  d  Overall deposition velocity [m/s]  xvi  Acknowledgements First and foremost I would like to thank my supervisors, Dr. Steven Rogak and Dr. Akram Alfantazi for their invaluable insight and guidance into making this thesis possible. I would also like to express my gratitude to Dr. Edouard Asselin whose advice gave further direction to this thesis. Furthermore, I would like to thank NSERC and AECL for their financial support in this research.  I would like to thank Andrej Boskovics for his help on the construction of the electrical heaters and flow loop. Additionally, I would like to thank my lab colleagues, Tirdad Nickchi, Hamid Zebardast, Hugo Tjong, Arka Soewono, Mohammad Taghavi, and Kamran Alba for their help, advice and support.  Also, many thanks to the technicians who were incredibly helpful and patient in teaching me the operation and analysis of the equipment: Dr. Berend Wassink on atomic absorption spectroscopy, Jacob Kabel on the scanning electron microscope, Anita Lam for X-ray diffraction analysis, the mechanical machine shop for their assistance in machining components, the electronics shop for their advice on the electrical heater and controls, and Perry Yabuno for always making sure that parts were ordered and received as soon as possible.  Finally, I would like to thank all my friends and family who provided moral support throughout the entire journey.  xvii  1 Introduction 1.1 Supercritical Water Supercritical water (SCW) refers to water above the critical point of 374⁰C and 22.1 MPa. Above this temperature and pressure, water exists only as a single phase fluid with very unique thermodynamic and fluid properties (Weingartner & Franck, 2005).  Supercritical water Pressure  Liquid 374ºC 22.1MPa  Solid Vapor  Temperature  Figure 1-1: Pressure - temperature diagram for water, grey region indicating supercritical conditions.  Viscosity, density, and dissolution of non-polar and polar species can change dramatically from subcritical to supercritical conditions. At subcritical temperatures, water is highly polar and is an excellent solvent for other polar substances such as inorganic salts. On the other hand, organic species are generally non-polar and therefore are usually not soluble in subcritical water. In supercritical conditions, the characteristics described above are reversed 1  and non-polar organic species and gases become significantly soluble, while ionic species become almost completely insoluble (Kritzer, 2004; Zhang et al., 2009).  While the critical temperature gives the threshold temperature for achieving supercritical conditions, when describing the temperature at which thermodynamic changes occur, it is often times more useful to describe the transition using the “pseudocritical” temperature. The pseudocritical temperature is the temperature at which fluid thermal expansion is at its maximum for any given supercritical pressure. This temperature increases as the pressure increases as shown in Figure 1-2. Other parameters such as heat capacity and ionic product of water are also affected by pressure and its transition also shifts to higher temperature as pressure is increased (Kritzer, 2004).  Figure 1-2: Density of water vs. temperature at various pressures, showing the effects of pressure on the pseudocritical temperature  2  1.2 Systems using Supercritical Water Medium Supercritical water has been a crucial medium in numerous industrial and research applications for its interesting properties. Large variations in fluid and thermal properties near the critical point offers an opportunity to fine tune fast chemical reactions with only small adjustments in pressure and temperature (Kritzer, 2004). The special properties of supercritical water as an excellent solvent for organic material have led to the technology of supercritical water oxidation systems (Shaw et al., 1991) while the excellent heat transport characteristics have led to power plants operating at higher temperature and pressure. Material degradation has been a major limiting factor, however, new technology in materials in the last few decades have allowed coal-fired power plants to operate in supercritical conditions (Viswanathan et al., 2005). Finally, the need for cleaner and more efficient energy has recently prompted research into supercritical water reactors (SCWR) for nuclear power plants. Despite the operational experience with supercritical water in various applications, corrosion and fouling in this medium remains a significant challenge.  1.3 Scope and Objectives Both the UBC Mechanical and Materials Engineering departments are collaborating with AECL in the development of monitoring techniques for corrosion and fouling in supercritical water. The primary application for the technology is for the CANDU supercritical water reactors, however, the scope of this research can be applied for many similar systems such as supercritical fossil-fueled and supercritical light water reactors. The purpose of this particular study is to explore the deposition and transport characteristics of magnetite formed in supercritical water with several experimental techniques and simulation models.  3  Experimental investigations were conducted to examine four techniques for measuring and monitoring fouling and transport of magnetite in supercritical water. These experiments were focused on determining the effects of heat flux, temperature and pH on the particle size, deposition rate, and adhesive strength of particle to the wall. A once-through flow loop was modified to conduct experiments of magnetite deposition in conditions which attempted to replicate SCWR conditions. A hydrothermal method for synthesizing iron oxide particles in supercritical water was adapted for producing simulated magnetite corrosion particles. Experimental procedures and description of the apparatus can be found in Chapter 4.  A high-current, electrically-heated test section with temperature control and tube-wall temperature measurements was constructed by the author for this study to measure fouling and removal rates. High and low temperature filters were also fabricated and installed after the test section to collect particles in each experiment. These filters were analyzed using Scanning Electron Microscopy (SEM) to provide insight into the effects of pH on the magnetite particle size and structure produced in supercritical water. In addition, deposition morphology and thickness was determined using SEM and Atomic Absorption Spectroscopy (AAS). A novel, two-step cleaning procedure was developed for qualitatively determining the bond strength of the magnetite deposit.  A heat and mass transfer model was created to compare simulations with experimental deposition thicknesses to determine if a mass transport limited model can accurately predict deposition of magnetite particles in supercritical water. Due to the lack of experimental data on the deposition mechanism of magnetite particles in supercritical water, deposition was  4  assumed to be mass transport limited. Simulations were also run to quantify the deposition of magnetite onto a CANDU nuclear reactor for one year. The results and its implications on a SCWR are presented in Chapter 6. Finally the conclusions and recommendations for future work are summarized in Chapters 7 and 8 respectively.  5  2 Literature Review 2.1 Applications of Supercritical Water 2.1.1 Supercritical Fossil-fueled Power Plants The thermal efficiency of fossil-fueled power plants has increased over the past century by operating steam at higher pressure and temperature, and thereby increasing the thermodynamic efficiency (Masuyama, 2001). In the 1920’s, the pressure and temperature was limited to 4 MPa and 370°C due to the mechanical properties of carbon steel. However, technological improvements in materials such as the development of nickel-based alloys and high temperature resistant steels have allowed coal-fired power plants to operate as high as 24 MPa and 600°C. With nearly two dozen power plants worldwide currently operating at 580 – 600°C and 24 – 25 MPa, there is considerable operational experience with supercritical water in power generation (Viswanathan et al., 2005). These experiences with supercritical fossil-fueled power plants can provide invaluable insight into possible operational issues for a nuclear SCWR.  Studies on the scaling of austenitic steels in supercritical fossil-fueled power plants show a duplex oxide structure in which the outer layer is likely to exfoliate compared to an inner layer of chromium and nickel oxide which almost never exfoliates. Once the scale thickness reaches approximately 100 – 200 µm, oxides on austenitic steel begin to exfoliate due to the differences in thermal expansion between the tube surface and the outer oxide layer, especially during startup and stop time of the boiler (Masuyama, 2001). It was also found that fine-grained steels and high-Cr steels had very slow oxide growth with very little  6  exfoliation. In general, chromium content above 20% was effective at preventing steam oxide buildup. Tube plugging due to exfoliated material buildup can cause overheating in the portion of the tube, resulting in creep rupture of the piping when the pressure due to the fouling buildup exceeds the creep strength of the tubing (Masuyama, 2001). In addition, exfoliated material has also been found to erode turbine components in a supercritical fossilfueled power plant if solid particles are found in the fluid stream. Therefore, scale build-up and its resistance to exfoliation is a major concern for any system running SCW.  There are several key differences between a supercritical fossil-fueled power plant and nuclear SCWR design requirements which make it particularly challenging for SCWR (Was et al., 2007). First, compared to supercritical fossil-fueled power plants which usually can have fire tubes with wall thickness of up to 10 mm, the fuel cladding in a nuclear reactor core is restricted to 0.5 - 0.6 mm due to thermal and neutron efficiency (Zhang et al., 2009). Larger amount of corrosion product deposits are often found on superheater tubes operating above the critical point in a supercritical fossil-fueled power plant and it is not unusual to find oxide films of several hundred micrometer thickness on these tubes (Cook & Fatoux, 2009). Such high fouling rates and deposit thickness would be unacceptable for a fuel cladding in a SCWR. Second, the cooling water in the boiler tubes sees a geometrically smooth surface along the boiler while the core of a SCWR will have a fuel assembly of intricate geometry (Was et al., 2007).  7  2.1.2 Supercritical Water Oxidation The high solubility of organic species in supercritical water has led to the research in “supercritical water oxidation” (SCWO), technology which was developed for destroying hazardous organic waste (Helling & Tester, 1988). Despite its initial success, problems with fouling and corrosion resulted in high operating costs which has forced the majority of commercial SCWO plants constructed and operated in the late 90’s and early 2000’s to permanently shut down (Marrone et al., 2005). There was extensive experimental work conducted on fouling and corrosion in supercritical conditions for SCWO applications, however, the environment is often highly oxidizing and in some cases highly acidic which are significantly different from the water chemistry which will likely be used for a CANDU SCWR (Kritzer, 2004). It has been well observed from SCWO research that acidic water chemistry promotes rapid corrosion, especially slightly below the pseudocritical temperature. With inorganic salts in supercritical conditions, very high fouling rates are observed and the deposit can completely plug flow of a system resulting in increased system pressure and the need for immediate shut down and cleaning (Teshima, 1997).  2.1.3 Nuclear Energy and the Generation IV Nuclear Reactors As energy demand continues to increase and the effects of energy production on the global environment escalates, the need for cleaner and more efficient technology grows ever more important. Nuclear energy provides approximately 13.8% of the world’s energy demand and its fuel provides energy with low greenhouse gas emissions making it an ideal candidate for future energy production (International Atomic Energy Agency, 2010). These facilities require some of the most advanced technology available and the Generation III++ is the most  8  recent design incorporating the technology developed for the industry over the last half century. Generation IV nuclear technology is to succeed the current Generation III++ nuclear reactors with state-of-the-art technology to improve on safety, lower byproduct waste, and higher efficiency (Cook & Fatoux, 2009; Khartabil, 2009). There are currently six emerging technologies for Generation IV nuclear reactors, one of these systems being SCWR in which the coolant will operate above the critical point of water. The first of the Generation IV nuclear reactors utilizing SCWR technology are expected to go online in 2025 as presented in Table 2-1 (Generation IV International forum SCWR Committee, 2002) .  Table 2-1: Generation IV systems and expected deployment year Generation IV R&D Systems Gas-Cooled Fast Reactor System Lead-Cooled Fast Reactor System Molten-Salt Reactor System Sodium-Cooled Fast Reactor System Supercritical-Water-Cooled Reactor System Very-High Temperature Reactor System  Year 2025 2025 2025 2015 2025 2020  The higher operating temperature and pressure of a SCWR is expected to increase the thermodynamic efficiency of nuclear power plants to 45 - 50% in comparison to the 33% 35% for conventional nuclear power plants (Naidin et al., 2009). This is due to the single phase coolant which improves heat transfer and simplifies the overall design by reducing the number of equipment and flow loops. In addition, the higher enthalpy of the coolant at higher temperatures especially near the critical point offers an opportunity to decrease the mass flow of the coolant for the same thermal output, resulting in smaller pumps, piping and housing (Generation IV International forum SCWR Committee, 2002).  9  2.1.4 CANDU Supercritical Water Reactors The CANDU (Canadian Deuterium Uranium) nuclear reactor is a Canadian-designed nuclear reactor developed by Atomic Energy of Canada Limited (AECL). The CANDU nuclear design is unique in that it uses heavy water as a moderator versus light water, which allows the reactors to use non-enriched uranium as fuel. In addition, the CANDU reactors use pressure tubes rather than a pressure vessel in the reactor core, simplifying high pressure designs and making the movement towards a supercritical coolant the next logical step (Torgerson et al., 2006).  The coolant would enter the reactor core at 350⁰C and leave the core at above 600⁰C and 25MPa, finally leading to a high pressure turbine in the case of a direct cycle system (Naidin et al., 2009). A simplified drawing of the proposed CANDU SCWR is provided in Figure 2-1.  The low solubility of inorganic salts in supercritical water suggest that there is greater possibility for higher fouling rates of metal oxides in a SCWR core than previously seen in a subcritical CANDU reactor. Metal oxides may nucleate from solution to form nano to micrometer sized particles which may foul onto various heat transfer surfaces, leading to higher than expected temperatures and accelerating material degradation. Particles which do not deposit may be carried in the bulk fluid to the steam turbines which can lead to solid particle erosion and out-of-core radiation (Guzonas et al., 2009).  10  Turbine  Generator  Reactor Core  Condenser  Preheater Pump  Figure 2-1:  Simplified drawing of a direct cycle coolant for Generation IV CANDU SCWR  2.2 Materials for SCWR There are currently five materials that are under evaluation as potential materials to be used in SCWR for various components such as the fuel cladding, pressure tubes, etc. (Zhang et al., 2009).   Nickel Alloys: 690, 625, 718    Austenitic Stainless Steels: 304NG, 316L, D9, AL-6XN    Ferretic / Martensitic Steels: P92, P112    Oxide Dispersed Steels (ODS): 9Cr-ODS, 12Cr-ODS, MA956    Modified zircaloys  11  Experiments presented in this study are constructed from mainly SS316L material. Austenitic Stainless Steels such as AISI 304 and AISI 316 are part of a family of stainless steels with nickel content which preserve the austenite crystal structure during cooling. Compared to martenitic and ferritic stainless steels, austenitic stainless steel enjoys the advantages of being easily formable, weldable, and ductile.  The ‘L’ of 316L indicates that the particular stainless steel has lower carbon content than that of unstabilized austenitic steels such as 316 and 304. When types 316 and 304, are subject to high temperatures in the range of 550°C to 880ºC, chromium carbide tend to precipitate along the grain boundaries, forming areas depleted of chromium below the minimum 12%. These areas become susceptible to corrosion in and near these boundaries where the material has become sensitized (Sinha, 1989). In operational power plants, these sensitized austenitic steels have been found to fail by intergranular stress corrosion cracking (IGSCC) in the tubing of boiler water reactors (R. L. Jones et al., 1993). By lowering the carbon content to below 0.03%, the amount of M23C6 that forms is minimized. Furthermore, the addition of molybdenum (Mo) in 316L SS as opposed to 304L SS which does not have Mo, improves corrosion resistance of stainless steel in chloride aggressive environments by forming a uniform passive layer (Bastidas et al., 2002).  Table 2-2: Stainless steel AISI 316L composition (American Society for Metals, 1985) C  Mn  Si  0.03  2.00  1.00  Cr 16.0 18.0  Ni 10.0 14.0  P  S  Mo  N  0.045  0.03  2.0 - 3.0  0.10  12  2.3 Corrosion and Fouling in Subcritical and Supercritical Water 2.3.1 Properties of Magnetite Magnetite is an iron oxide mineral with the chemical composition of Fe3O4. It is also known as a ferric-ferrous oxide as it is one part wüstite (FeO) and one part hematite (Fe2O3). It is part of the spinel group which is a group of minerals that have a chemical composition of A2+B23+O42-. It typically forms as an octahedral, less commonly as dodecahedral and rarely as a cubic (Anthony et al., 1997). Magnetite, which is the stable form of iron oxide in SCWR conditions, is expected to be the dominant metal oxide formed in the core due to the presence of iron in many high temperature alloys. It is a common corrosion product found on the tubes of fossil-fueled and nuclear power plants (Propp et al., 1996).  Pourbaix diagrams, also known as Eh-pH diagrams indicate the stable regions for metal species. Solid lines indicate the border between two species while the diagonal dashed lines indicate the stable region for water. Above the upper dashed line, water is oxidized to oxygen while below the lower line, water is reduced to hydrogen (D. A. Jones, 1996). The equilibrium equations of water in neutral or alkaline solutions are shown below:  2-1  2-2  Figure 2-2 is a Pourbaix diagram for iron and iron-chromium created from SUPCRT92. Because of its dependence on temperature and pH, the borders for stable species will vary  13  from one temperature to another and change dramatically in supercritical conditions (Propp et al., 1996). 2.20 Fe3+  1.45  FeO42FeO42-  0.70 Fe2O3  Fe2O3  2+  Fe  -0.05  2+  Fe Fe3O4  Fe3O4  -0.88 Fe  Fe  Eh(V)  -1.55  HFeO2  2.20 Cr2O72-  1.45  FeO42-  CrO42-  CrO2  0.70  Cr2O72-  FeO42-  CrO2  CrO42-  3+  Cr  Fe2O3  -0.05 Fe2+  -0.88  Cr3+  Fe2O3  Cr2O3  Cr2+  Cr2O3  CrO Fe  -1.55  Fe  Cr  CrO33Cr  -2  2  6  10  14  -2  2  6  10  14  pH  Figure 2-2: Pourbaix diagram (a) Fe – H2O at 25ºC and 1 atm, (b) Fe – H2O at 374°C and 220 atm (c) Fe/Cr – H2O at 25°C and 1 atm, (d) Fe/Cr – H2O at 374ºC and 220 atm. Based on SUPCRT92 (Propp et al., 1996)  14  2.3.2 Magnetite and its Solubility in High Temperature and Pressure Water The solubility of corrosion products play a primary role in corrosion and fouling in aqueous systems (Kritzer, 2004). In particular, the solubility of magnetite, an iron oxide of the form Fe3O4, in high temperature and pressure water is a crucial information to operators in the power industry due to the prevalence of the mineral found on heated tube surfaces.  Solubility of magnetite in alkaline and acidic water conditions has been measured up to 300⁰C by several groups. Sweeton and Baes Jr. (1970) tested magnetite solubility with varying pH using potassium hydroxide and hydrochloric acid while Tremaine and LeBlanc (1980) adjusted pH with sodium hydroxide and hydrochloric acid. The dissolution of magnetite requires the reduction of FeIII to FeII. The equilibrium equation is provided in equation 2-3 (Tremaine & LeBlanc, 1980).  ( )  ( )  (  )  2-3  In order to predict the magnitude of super-saturation and potential fouling rates in supercritical water, the solubility of magnetite under these conditions is required. Unfortunately, the experimental solubility data for magnetite in supercritical water is currently unavailable. There are several major problems which make measuring solubility of magnetite extremely difficult. First, the low solubility of iron means that measurements must be done in parts per billion and parts per trillion ranges. Preventing iron contamination of the solution from tubing, pumps, valves, etc. becomes increasingly difficult as materials that withstand supercritical conditions are limited to high temperature alloys which often have iron content eg. 316L, Inconel 625, Alloy 800.  15  To overcome the deficiency in experimental solubility data for magnetite, thermodynamic studies utilizing the Helgeson-Kirkham-Flowers (HKF) extrapolation for the Gibbs free energy at high temperature and pressure have been used to predict the solubility of magnetite in supercritical conditions (Cook & Fatoux, 2009; Guzonas et al., 2009). In another study by Burrill (2000), magnetite solubility data from Bohnsack was extrapolated from subcritical to supercritical conditions, assuming a direct dependence of magnetite solubility with water density. The simulation results are provided for comparison in Figure 2-3.  Figure 2-3: Magnetite solubility adapted from Cook and Fatoux (2009) for neutral pH water at 25 MPa, Guzonas et al. (2009) for pH = 9.3 at 30 MPa, Burrill (2000) for neutral pH at 25 MPa, and Sweeton and Baes Jr. (1970) experimental results.  16  All three solubility simulations show a drop in solubility between 350ºC to 400ºC as the density of water drops and the water transitions from polar to non-polar solvent. For models based on HKF, the solubility increases once again as the temperature increases even higher after the initial drop. The HKF model has been successful in correlating equilibrium constants for hundreds of inorganic aqueous species in high temperature and high pressure water including supercritical water (Sue et al., 2002). Cook and Fatoux model was adopted for the simulation in this study for this reason as well as because it used pressure closer to those expected in a CANDU SCWR of 25 MPa. Under this model, minimum solubility is reached at 400°C and 25 MPa under neutral pH conditions with an iron concentration of approximately 0.2µg/L. At supercritical conditions, neutral species such as Fe(OH)2 and Fe(OH)3 are expected to be the dominant dissolved species in the solution (Cook & Fatoux, 2009).  2.3.3 Corrosion Due to significant changes in the thermal and chemical properties of water from subcritical to supercritical conditions, corrosion rates and mechanisms can be substantially different in supercritical water environment (Was et al., 2006). In supercritical water systems, general corrosion and stress corrosion cracking are found to be one of the dominant corrosion mechanisms (Was et al., 2007; Zhang et al., 2009). For general corrosion, the surface of the alloy undergoes oxidation and forms metal oxide corrosion product. This can be of particular concern since SCW can be an aggressively oxidizing environment if certain species or dissolved oxygen are present, resulting in very high corrosion rates (Lister, 1980). It has been  17  found experimentally that at high temperatures, oxides are preferentially formed over hydroxides as the primary form of corrosion product (Kritzer, 2004).  However, not all corrosion products are undesirable and the corrosion products can produce a protective film that reduce the rate of corrosion of the underlying metal (Lister, 1980). Corrosion of austenitic stainless steel and nickel alloys subject to supercritical water is observed to form a multilayer oxide on the surface. These are found to be tens to hundreds of nanometers thick after several hundred hours in a deaerated supercritical condition (Was et al., 2007). In high temperature aqueous systems, iron ions diffuse to the outer surface while oxygen from the solution diffuses into the metal alloy. This produces a duplex oxide structure: a dense inner chromium-iron rich layer and a porous magnetite outer layer. The inner chromium rich layer has a spinel structure and provides protection to the alloy while the porous outer layer provides no further protection (Gao et al., 2007; Was et al., 2006). Therefore, the corrosion rate is dependent on the oxygen diffusion through the inner layer following an Arrhenius behavior in which the high temperature provides the activation energy for diffusion to occur (Zhang et al., 2009)  2.3.4 Fouling Another method for corrosion products to form on the surface is the deposition of corrosion products from the bulk solution through “ex-situ” fouling. Ex-situ fouling can be further divided into two subcategories of crystallization and particulate fouling. In crystallization fouling, the bulk fluid is super-saturated with, for example, iron species which may have come from other parts of the power plant where corrosion had occurred and been transported  18  to the surface of interest. Once these iron species reach the wall, super-saturation causes particles to nucleate and deposit onto the wall. In particulate fouling, nucleation happens in the bulk fluid and the corrosion products are carried to the wall as solid particles. Lister (1980) hypothesized that the deposits of these ex-situ corrosion products may be the same form as the outer layer oxide formed from the alloy. As the iron species diffuse from the alloy to the surface, the iron first dissolves into the solution. Then, due to local supersaturation, it re-deposits as porous outer layer resembling ex-situ fouling.  Ex-situ fouling of oxides onto heat transfer surfaces can be detrimental to power systems as they can decrease thermal efficiency and increase the temperature of the heat transfer material promoting earlier degradation (Lister, 1980). To track thermal efficiency issues due to deposition, heat transfer performance is regularly monitored on steam generators in pressurized water reactors (PWR) (Schwarz, 2001). The steam generator of a PWR is a heat exchanger where the reactor coolant transfers the heat to the feed water / steam cycle. The accumulation of deposits, mainly iron hydroxides and oxides was found to impair heat transfer between the two systems. Schwarz (2001) compiled data from twelve Siemens steam generators, many spanning over a decade and one for 27 years. Fouling was determined using energy balance and heat exchanger equation shown below.  2-4  where Q is the thermal energy into the system, ΔTlog is the mean logarithmic temperature and U is the overall heat transfer coefficient for the heat exchanger, calculated by equation 2.3.  19  2-5  where Rf is the fouling factor representing the additional thermal resistance due to the iron oxide deposit. The thermal power Q was determined from the change in enthalpy of the feedwater, derived by the inlet and outlet temperatures at constant pressure. Accuracy of the fouling measurements was found to be highly sensitive on the accuracy of the steam generator outlet temperature and pressure measurements. Using the heat exchanger and the calculations described above, it was possible to observe the fouling rates on the steam generator over its operational time. It was found that switching from phosphate-treatment to an all-volatile treatment (AVT) with hydrazine to ensure high pH, significantly reduced the fouling rate on the SG. The fouling factor increased gradually over time under phosphatetreatment which occasionally would encounter a sharp drop after undergoing chemical cleaning (Schwarz, 2001).  Experimental, simulations and theoretical studies of magnetite deposition in aqueous systems has been the subject of many previous studies. A classic study of particle deposition was conducted by Thomas and Grigull (1974) using magnetite powder in high temperature and pressure system. Their experiment used an electrical heater in a flow loop system and is one of the few studies which reached supercritical conditions of 24.6 MPa and 471°C. They observed that the deposition of magnetite in both subcritical and supercritical single phase flow was approximately linear to Reynolds number and concentration which is an indication that the deposition was mass transport limited (Thomas & Grigull, 1974). More recent studies by Turner and Klimas (2000) and Basset et al. (2000) used sol-gel methods to form  20  narrow size-distribution magnetite particles for deposition experiments. Deposition experiments have shown the effects and differences of deposition of single phase, boiling, and non-boiling two phase flow. Under single-phase flow, removal was not found to be a significant factor and only significant in sub-cooled boiling (Basset et al., 2000).  Since the proposal of the CANDU SCWR reactor design, deposition of corrosion products in the reactor core has been identified as a problem by a number of researchers prompting simulations for deposition to be developed. Using the solubility mentioned in section 2.2, Burrill (2000) created a simulation with a CANDU 6 fuel channel design and the temperature went from 350°C to 384°C at 25 MPa. Ferrous ions were assumed to diffuse towards the wall and crystallize as magnetite deposits. Deposition was assumed to be mass transport limited and the diffusivity coefficient of ferrous species was calculated using empirical formulations by Miller (1982). Burrill (2000) determined that after one year of operation the peak deposit was 108mg Fe/cm2 occurring 4m from the channel inlet. The effect of dissolved iron concentration on deposition was also investigated by setting the concentration at the inlet of the reactor core to 18.22 µg/L and 5 µg/L. The higher inlet concentration led to deposition nearly six times higher than the lower concentration, suggesting that the uncertainty in magnetite solubility can have a significant impact on the simulation results (Burrill, 2000).  In a similar study, Cook and Fatoux (2009) proposed the use of steam generator such as the one used in a PWR for the CANDU SCWR to prevent corrosion products from travelling to the steam turbine. A simulation using the HKF model for saturation of iron was used and the kinetics of mass transport was arbitrarily assumed to be described by constant deposition  21  velocities of 0.01 cm/s and 0.001 cm/s. The peak deposition thickness in the steam generator was calculated to be up to 300µm or 50µm depending on the kinetic constant assumed. The deposit decreased the efficiency of the steam generator by 4% after five years of operation (Cook & Fatoux, 2009).  2.3.5 Corrosion Product Removal and Transport Once attached to the tube surface, adhesive forces prevent the particle from flowing back into the fluid stream. Adhesive forces are generally classified into three categories; The first class of adhesive forces is intermolecular forces which can include van der Waals, the second are chemical bonds including hydrogen bonds, and the third is sintering effects such as diffusion. (Cooper et al., 2001).  As particles deposit, they are located in the viscous sub-layer, a very thin layer near the wall where viscous effects are dominant and the flow is laminar (Cengel & Cimbala, 2006). Detailed investigations of particulate removal by Cleaver and Yates (1973) have shown that the viscous sub-layer is unsteady in turbulent flows, and that there are turbulent “bursts” that can cause particle removal when the shear flow generate lift forces exceeding adhesive forces of the particle and surface. In contrast, if the viscous sub-layer thickness is much greater than the particle deposited, the viscous forces can rapidly dissipate the turbulent bursts (Cengel & Cimbala, 2006; Turner et al., 1990). Previous studies for calculating removal rates have used an exponential relationship (equation 2-6) and incorporated the frictional velocity U* (equation 2-7).  22  ̇  The value  ̇  ̇  2-6  is found experimentally, t is time [s] and λ is removal rate constant [1/s]  calculated using equation 2-7.  2-7  where f is the friction factor and ν is the kinematic viscosity.  While fouling of corrosion product on heated surfaces is of concern, the transport of particulates outside the reactor core is also of great importance to the reliability and safe operation on a power system.  Solid particle erosion (SPE) can be problematic in power generating systems where the protective oxide layer of piping and other components are eroded by particles suspended in the fluid stream (Crockett & Horowitz, 2010). SPE can contribute to increased maintenance costs and the decrease in the thermal efficiency of a turbine which can be expensive (Dai et al., 2007). For a SCWR, a direct cycle from the reactor core to the steam turbine is expected and the problem of SPE may be enhanced due to lower solubility of oxides in supercritical water and the formation of particulates. Another issue with corrosion products being transported in a nuclear reactor is the transport of radioactive material. Corrosion products can form in the reactor core, attaching to fuel cladding and pressure tubes which are then irradiated (Burrill, 1977). Once the particle exfoliates, it can travel to other parts of the  23  system outside the core and become a source of hazardous radiation to its operators and maintainers (Lister, 1980).  The need for in-situ particulate sampling systems can be significant for power plant applications. Information about the properties of the particle such as size, shape, hardness and strength can lead to information about potential damage that particulates can have on system components surfaces (McCabe et al., 1985). The composition of the particles give an indication as to where they may be coming from (eg. fuel cladding, piping, etc.) as well as a predicted hardness and strength of the particles.  Furthermore, sampling at high temperature has been shown to provide better representation of the particulates in the system. Turner and Klimas (2000) made measurements of particulates before and after a filter and found that by placing the filter before the heat exchanger, the concentration of iron was twice as large as when the filter was sampled cold. Turner argued that such difference is likely due to thermophoresis where the particles diffuse down a temperature gradient, therefore transported to the colder surface and attaching to the heat exchanger wall. In addition, sampling ports in real power-generating systems are often far away from the sampling line which can cause losses in the particulates found at the ports. In PWR, it was found that 50-75% of particles were lost or dissolved in the solution due to long sampling lines (Srisukvatananan et al., 2007)  24  3 Heat Transfer and Deposition Modeling 3.1 Overview A heat transfer and deposition model was created in MATLAB to aid in the design of the experiments and to predict the effects of magnetite deposition on a CANDU SCWR fuel cladding. Only particle deposition was considered in this model due to the lack of data on the deposition rate constant for ferrous species in supercritical water. Simulation models were also used to compare expected deposition thicknesses with those found in the experiments. The heat transfer and deposition model for the CANDU SCWR and the experimental simulation are based from the same code. However, due to different geometry and flow characteristics, adjustments to the code were made where necessary. The MATLAB program source code can be found in Appendix A.  3.2 Heat Transfer Calculations of Water 3.2.1 Enthalpy of Fluid in the Test Section The test section in the once-through flow loop is electrically heated and the heat flux is assumed to be constant throughout the test section. As a result, the enthalpy of the bulk fluid is a linear function of the distance along the tube, and at any point, the enthalpy can be determined from equation 3-1:  ̇  3-1  25  where H1 and H2 is the enthalpy [kJ/kg] at a distance Δx [m] away, ̇ is the mass flow rate [kg/s], and q is the heat input per meter [kW/m]. The enthalpy of the inlet and outlet of the test section is determined by using the inlet and outlet bulk fluid temperatures [⁰C] at the specified pressure P [bar]. The bulk fluid temperatures were measured experimentally using in-situ thermocouples before and after the test section, as described in detail in Chapter 4.  3.2.2 Heat Transfer Coefficient in Supercritical Water Heat transfer correlations developed for subcritical conditions are not applicable for supercritical conditions due to the dramatic changes in the thermodynamic and transport properties near the critical region. Therefore, a number of correlations have been developed for supercritical conditions in previous studies. Calculating non-dimensional numbers for heat transfer generally require the use of either bulk or wall temperatures, however, there can be dramatic changes in the heat transfer properties if bulk and wall temperatures are near the pseudocritical region. Correlations such as those developed by Swenson et al. (1965) and Yamagata et al. (1971) attempt to reflect these changes in the properties by using both wall and bulk temperature to determine a suitable average. The equations provided by Swenson are a relatively straightforward method for calculating the Nusselt number and has been shown to predict heat transfer well, demonstrated by Teshima (1997).  (  )  3-2  26  By using equation 3-2, one can determine the Nusselt number for finding the heat transfer of the tube to fluid. This requires the Reynolds number as well as the Prandtl number found in equation 3-3. ̅  3-3  where ̅ is the integrated average heat capacity.  ̅  ∫  3-4  A recent study by Bazargan and Fraser (2009) on the heat transfer correlation in supercritical water in a horizontal pipe showed that many previous correlation used were insufficient in predicting the Nusselt number accurately. Bazargan concluded that large variations in fluid properties near the pseudocritical zone dramatically enhanced heat transfer. A comprehensive explanation and discussion of the approach is described by Bazargan and Fraser (2009).  Both the correlation of (Swenson et al., 1965) and (Bazargan & Fraser, 2009) were programmed in MATLAB for comparison between the empirical models and experimental results. These results are presented in section 4.5.1. The Nusselt number calculated from the correlations is then used to calculate the wall temperature using equation 3-5.  3-5  To determine the fluid properties at the wall, the wall temperature is required in the previous calculation. Since this is the parameter which is currently unknown, a wall temperature must  27  first be predicted by the software. The program runs through an iterative loop, first calculating the wall temperature with the initial guess, calculating the new wall temperature and modifying the initial guess until the two temperature converges to within 0.1ºC.  3.2.3 Temperature of the Oxide Layer As the particles begin to foul the inner wall of the tube, the deposits create a heat transfer resistance between the fluid and tubing wall interface. The temperature at the wall calculated from section 3.2.2 is therefore the temperature of the oxide/fluid layer as deposition occurs. The temperature of the oxide/tube layer can be calculated as follows:  3-6  where koxide is the thermal conductivity of the fouling oxide (Holman, 2002). Magnetite thermal conductivity was calculated using the following empirical formula  3-7  where T is the temperature of magnetite (Electric Power Research Institute Inc., 2003).  3.2.4 Temperature of the Outside Wall Temperature Calculating the outside wall temperature of the tube was necessary because the thermocouples on the test section are measuring outside wall temperature. The tube has internally generated heat from the low AC voltage and high current applied to the tube. The  28  equation for determining the external wall temperature with internally generated heat assuming a perfectly insulated outside wall is shown in equation 3-8.  (  )  3-8  where A is the ratio of the inner and outer radius found by equation 3-9.  3-9  3.2.5 Buoyancy Calculations Buoyancy effects were calculated for the experimental conditions to determine if it would have an effect on the temperature profile of the tubing wall. Buoyancy effects can be significant for fluids near the pseudocritical point due to large variations in the fluid density at the bulk and wall temperatures (Bazargan et al., 2005). This can cause the low density supercritical water to flow at the top of the tube while the denser subcritical water flows at the bottom of the tube. Since the heat transfer coefficient varies dramatically from supercritical to subcritical, this means that a difference in top and bottom surface temperatures can have significant implications for a horizontal flow in tube.  Buoyancy effects were calculated according to Petukhov et al. by a threshold value for the Grashof number, which below the threshold the buoyancy can be neglected. The threshold is defined as follows:  29  ̅̅̅  ̅̅̅  3-10  where  ̅̅̅  3-11  The Grashof number is calculated using a heat-flux-related definition of equation 3-12. ̅ 3-12  where  ̅  3-13  Evaluation is done by determining if Grq < Grth or that the Grq/Grth is less than one.  30  Figure 3-1: Calculation of Grq/Grth for determination of the effects of buoyancy along the test section using Bazargan et al. (2005)  Since Grq<Grth for all experiments run, buoyancy should not affect the temperature of the outer wall.  3.3 Deposition Calculations Fouling is a complex phenomenon in which many of the parameters are unknown or specific to the condition of deposition. In a real SCWR core, formation of magnetite may be a combination of particles formed in the bulk solution and crystallization of magnetite on the tube wall. In order to develop a practical simulation with limited data available, certain assumptions for the formation of magnetite and deposition characteristics are made and presented. For the simulation under experimental conditions, particles are assumed to be of uniform size and the formation is assumed instantaneous at the inlet of the test section. For  31  the CANDU SCWR simulation, the particles are assumed to form instantaneously after super-saturation.  In addition, corrosion products which foul onto the heat transfer surface can come from the tube surface itself or from the environment. For the case of the experimental simulation, only ex-situ fouling will be considered since the experiments were run for a short period (40 minutes) making corrosion negligible. For CANDU SCWR simulation, corrosion is again neglected because it will be highly dependent on the material used for the fuel cladding which is currently undecided.  There are five sequential steps in fouling; induction, transport, attachment, aging and finally removal (Turner, 1993). For particulate fouling, the induction period is not important and is primarily limited by transport and/or attachment. Early deposition models tended to use a single mass flux term which depended on the concentration. A sticking probability ‘S’ was used to take into consideration the probability of particles reaching the wall without sticking, depending on the conditions of the particle of interest. In later models, the sticking probability was replaced with equations of surface attachment. This is described in the Kern and Seaton model which is most commonly used in particulate fouling (Epstein, 1987). The deposition is said to be a two-step process; mass transport where the particle moves from the bulk to near the surface, and an attachment process where the particle attaches to the wall. Therefore, the deposition is the net resistance of the two processes and can be described by the equation below.  32  3-14  where Kd, Kt and Ka is the effective, transport and attachment deposition velocity [m/s] respectively (Basset et al., 2000).  The term  is the attachment deposition velocity and is believed to be caused by van der  Waals interactions as well as the electrical double layer when the particle is approximately 10nm from the surface (Turner, 1993). Van der Waals forces arise from dipole-dipole and dipole-induced-dipole interactions between particle and wall (Turner & Klimas, 2000). These forces are always attractive between a particle and wall in a water medium (Hong, 2007). The electric double layer forms at the surface of an object that is immersed in a liquid. Solid objects generally attain a surface charge when immersed in a liquid which can be formed through the absorption of ions present in the liquid or the dissociation of molecules on the surface of the object (Bott, 1995). This creates a redistribution of ions in the medium in near vicinity to the immersed particle with similar charged ions being repelled and opposite charged ions being attracted. Therefore, a second layer in the medium called the diffuse layer is formed by coulomb forces with oppositely charged ions. As a particle approaches the wall, the diffuse layers of the particle and wall in the solution overlap and create a potential energy. If the charges are oppositely charged, the forces are attractive and the limiting rate is the mass-transport velocity. If the charges are the same, the force is repulsive and depending on the magnitude of the repulsive force, the overall deposition may be surface attachment limited (Basset et al., 2000). In such case, the particle that approaches the wall will have to  33  overcome this potential energy barrier to attach to the wall. This follows an Arrhenius behavior and can be described as: ⁄  3-15  where E is the activation energy, R is the gas constant, A is an experimentally determined factor and Ts is the temperature of the surface (Turner, 1993).  When deposition is mass transfer controlled, Ka is much larger than Kt and the mass transport velocity is approximately equal to the effective deposition velocity.  3-16  For deposition onto the inside of a tube, the mass deposition rate along any section of the tube  is then determined by: ̇  where  3-17  is the density of the particle fouling the surface. The deposition velocity is  calculated from equation 3-18.  3-18  where U* = frictional velocity and determined by:  34  (  where  is the fluid density and  )  ⁄  3-19  is the wall shear stress and calculated by:  3-20  The friction factor f in a fully developed turbulent pipe flow can be calculated using the wellknown Colebrook equation. In many cases an approximate but explicit relation can be substituted for the Colebrook friction equation using Haaland’s equation shown in the equation below.  (  where  (  )  )  3-21  is the roughness factor which is was taken to be 0.002 mm for a clean stainless steel  tube. The uncertainty in such a roughness factor can be as much as ±60% (Cengel & Cimbala, 2006). As the tube begins to foul onto the surface, it is assumed that the roughness of the tube was then the sum of the deposit and initial roughness:  3-22  The non-dimensional deposition velocity Kd+ was predicted from empirical equations fit by Papavergos and Hedley (1984) for aerosol particles landing onto a flat plane. Experimental work done by Teshima and Khan showed that these equations provided agreement for salt deposition in supercritical water (Khan, 2005; Teshima, 1997). 35  ( )  3-23  { Kd+ depends on both the dimensionless Schmidt number as well as heavily on the particle relaxation time. The Schmidt number is given in equation 3-24.  3-24  where Sc = dimensionless Schmidt number µ = dynamic viscosity of the fluid [Pa∙s] D = diffusion coefficient [m2/s]  The diffusion coefficient was determined from the Stokes-Einstein equation:  3-25  where bk = Boltzman’s constant [J/K] µw = dynamic viscosity of fluid [Pa∙s] Tw = temperature of fluid [⁰C]  Finally, the non-dimensional parameter relaxation time is calculated using the following formula (Turner, 1993):  36  (  )  3-26  The regions of Kd+ can be categorized into three regions depending on the relaxation particle time. For  , the particle motion is dominated by Brownian diffusion and deposition is  inversely proportional to the size of the diffusing particle (Beal, 1970). As the particle size increases, the relaxation particle time increases and momentum effects gain importance. The deposition velocity reaches a minimum when  is 0.2 and inertial coasting becomes the  dominant mechanism for transport which increases exponentially with diameter size. At even higher relaxation particle times above 20, the particle inertia becomes very large and the stopping distance exceeds the viscous and buffer layers. The stopping distance marks the distance at which the particle will coast to the wall simply by their momentum. In this region, the particles are no longer affected by the fluid forces, therefore reaching a constant value (Chen & Ahmadi, 1997).  Figure 3-2 was produced using the equations described above and shows the effects of particle diameter and temperature on deposition velocity under the experimental conditions.  37  Figure 3-2: Effect of particle diameter and fluid temperature on deposition velocity for experimental apparatus  3.4 CANDU Supercritical Water Reactor Simulation To predict the rate of deposition that might occur in a CANDU SCWR, the design of a CANDU ACR-700 was used as a baseline since the design for a SCWR pressure tube has yet to be finalized. The CANFLEX 43-rod fuel bundle configuration which is used in the ACR700 provided geometric parameters for the heat transfer surfaces and is listed Table 3-1.  38  Table 3-1: SCWR and fuel assembly design parameters Coolant pressure (MPa) (Naidin et al., 2009) Inlet temperature (⁰C) (Naidin et al., 2009) Outlet temperature (⁰C) (Naidin et al., 2009) Maximum cladding temperature (⁰C) (Naidin et al., 2009) Coolant total flow rate (kg/s) (Naidin et al., 2009) Number of fuel channels (Naidin et al., 2009) Heat flux (kW/m2) (Li et al., 2009) Fuel elements diameter (mm) (Li et al., 2009) Fuel elements length (m) (Li et al., 2009) Pressure tube diameter (m) (Li et al., 2009)  25 350 625 850 1320 300 1000 12 6.0 0.10  As an approximation of a real fuel bundle which has complex geometry, the simulation developed in this study was conducted by looking at a single fuel element and approximating the fluid and thermal properties around the element without the influence of neighboring elements. Temperature rise of the coolant was predicted by setting the fuel cladding wall as the single source of heat to increase the enthalpy of the coolant linearly along the 6 m distance.  For the hypothetical single fuel element developed in this study, the fuel bundle was converted into an equivalent heated tube with the coolant flowing on the inside, and therefore the MATLAB code used for this simulation is nearly identical to the simulation used for the experiment with the exception of the parameters. The mass flow rate for a single fuel element was calculated by the total flow rate in the reactor divided by the 43 fuel elements. An effective diameter for a single fuel element was calculated by keeping the velocity of the fluid the same as the velocity in a full scale fuel bundle.  39  Pressure Tube  Cladding Coolant Tbulk Toxide Tcladding  CANFLEX fuel bundle  Equivalent tube  Figure 3-3: Cross-sectional view of CANFLEX 43 rod fuel bundle and drawing of equivalent tube for simulation calculations (not to scale)  The surface most interesting for these deposition studies is on the fuel cladding which covers the fuel pellets and prevents the uranium from coming into contact and contaminating the water. The fuel cladding is often made zirconium alloys due to its low neutron absorption property compared to other alloys. However, zirconium alloys are limited by their reduced mechanical properties at high temperature and therefore may require other alloys as a substitute for the fuel cladding for future SCWR. Results from the simulations are reported in Chapter 6 and discussed considering the measurements described in the next few chapters.  40  4 Experiments 4.1 Supercritical Water Once-through Flow Apparatus A once-through flow apparatus shown in the Figure 4-1 was operated for all high temperature experiments reported in this study. The apparatus was originally constructed for running experiments for supercritical water oxidation and therefore, system components were added and/or modified to run the system for corrosion product transport experiments and simulate supercritical water reactor conditions. A schematic of the entire system can be found in Figure 4-2.  Back pressure regulator  Control Panel  Effluent tank Injection tank  Injection pump  Figure 4-1: SCW system used in all experiments located in the Mechanical Engineering laboratories.  41  Nitrogen  Nitrogen  Primary Tank  Relief Valve P-1  T2  T1 Preheater 1  Preheater 2 Nitrogen  Pump 1 Back Pressure Regulator  P-2  T-C  Tb1  Tb2 Primary Heat Exchanger  Test Section  Injection Tank  Preheater 3  LTHP Filter  Effluent Tank  Secondary Line Heat Exchanger Pump 2  Effluent Tank  LTLP Filter  HTHP Filter  Figure 4-2: Schematic of supercritical once-through flow apparatus 42  The majority of the system tubing is constructed from 1/4” seamless AISI stainless steel 316L. Tube fittings for the system use SS316 Swagelok fittings where contact with high temperature and pressure water is present. A 20L plastic tank and a glass flask were used for the primary feedwater tank and the injection solution tank respectively. It was important to keep oxygen content low in the system favoring reducing conditions, and therefore the tanks were deaerated by vigorously bubbling nitrogen into the water.  A high pressure Bran Leubre positive displacement pump was connected to the primary tank for delivering the majority of the fluid. Flow rates are manually controlled and can attain up to 8L/hr at a pressure of 25.0 MPa. To suppress the fluctuations in the pressure due to the positive displacement pump, a pulsation damper (Flowguard DS-10-NBR-A) pre-charged with nitrogen was installed near the Bran Leubre pump.  Three ceramic preheaters with a combined total wattage of 8.6kW are located downstream of the main pump which externally heat the fluid to the desired temperature before entering the test section. The ceramic heaters are split into top and bottom halves allowing for easy access to the tube during cooling and setup.  An Eldex Laboratories high pressure pump (BB-4-VS) injects the ferrous precursor solution into a flow-tee which joins the injection stream to the hot fluid stream. The pump has two heads, both taking fluid from the injection tank and reconnecting through a tee at the outlet. The pump has a manually adjustable flow rate using a Bodine Electric Company, RPM controller. To ensure complete thermal and chemical mixing of the two fluids, a perforated stainless steel 304 was  43  placed inside the tube immediately prior to the bulk thermocouple. A detailed description of the test section and the temperature measurements are given in section 4.7.  After the test section, a final tee splits the stream towards a secondary line consisting of a high temperature, high pressure filter and the primary line connecting directly into a heat exchanger. The secondary line has its own heat exchanger and a needle valve to cool and depressurize the fluid to room temperature and ambient pressure. A low temperature, low pressure filter is located after the needle valve in the secondary line.  On the primary line, a 2.0 m long tube and shell heat exchanger cools the hot supercritical fluid flowing in the tube side with cold water flowing in the shell side. After the primary heat exchanger, a low temperature high pressure filter and a back pressure regulator (TESCOM 261762-24) are located immediately downstream to adjust the pressure of the system  4.2 Flow and Temperature Control 4.2.1 Safety The safety of the operator, as well as other laboratory personnel was considered of utmost importance. Hence, online temperature and pressure monitoring with automatic shutdown control was put into place. Temperature was monitored at five locations throughout the system to keep temperatures below tube temperature ratings. If temperatures of the heated section were to exceed 500⁰C, the controllers are programmed to trip the breaker causing all heaters to turn off automatically while the primary pump remains online to remove residual heat from the system.  44  Pressure is also monitored by the controller and can turn off both the heaters and the pump if the pressure exceeds 34.5 MPa, signaling a possible plug in the system due to fouling. In addition, a pressure relief valve just downstream of the pump was set to open at 27 MPa.  If temperatures at the exit of the heat exchanger begin to rise, this may indicate the cooling water may be off or have insufficient flow rates which may result in high temperature water/steam exiting the system and potentially damaging downstream equipment. The controller was set to automatically turn off the heaters if the temperature of the tube after the heat exchanger exceeded 50⁰C.  4.2.2 Pressure Measurements Pressure was monitored with a GP-50 pressure transducer (P-1 in Figure 4-2), calibrated by the manufacturer with an accuracy of ±2% full scale output (FSO) of 34.5 MPa. These were read and recorded manually from the control panel board. A Bourdon pressure gauge (P-2 in Figure 4-2) downstream of the primary heat exchanger was also frequently checked as a safety measure.  4.2.3 Flow Rate Measurements The primary pump was first turned on and pressurized to the desired pressure. Flow rates of the effluent were measured three times and averaged using a graduated cylinder and a stop watch. Then the injection pump was turned on and after several minutes the flow rate was measured once again. The calculated difference between the two averaged flow rates was the injection pump flow rate.  45  4.2.4 Fluid Conductivity Measurements An Omega CDH-287 micro-conductivity meter was calibrated using a CDE3052 conductivity standard (1413µS/cm). The conductivity meter has an accuracy of ±0.3% FSO. Conductivity of the inlet of the primary tank was measured to check that the deionized water had low conductivity indicating ions are not present. The conductivity meter was also used for high pH experiments for measuring sodium hydroxide content in the tank. A detailed description of this calibration and use for high pH experiments is explained in section 5.5.3.  4.2.5 Dissolved Oxygen Concentration Measurements An Omega DOB21 measured the dissolved oxygen (DO) content in the two feedwater tanks. The DO has a relative accuracy of 0.02% FSO, with a minimum resolution of 0.01ppm.  4.3 Particle Synthesis for Modeling Power Plant Fouling Due to the low solubility of magnetite, deposition of magnetite from a dissolved state was expected to be slow, taking several years to form measurable deposit thicknesses. To create a feasible experiment, it was necessary to introduce ferrous species artificially. To simulate ferrous corrosion products, ferrous salts were mixed into deionized water and injected into the system. The ferrous salts then form iron oxide particles after the process described here.  Formation of magnetite particles has been of interest in colloidal science for its industrial applications, as well as through its presence in the power generating industry as corrosion products from iron and steel (Zhao et al., 2009). Work of Sugimoto and Matijevic (1980)  46  introduced a procedure of producing spherical magnetite particles using ferrous hydroxide gels. Later work by Adschiri et al. (1992) reported a rapid hydrothermal process using supercritical water to form sub-micron metal oxide particles in a continuous production. Hydrothermal processes for producing oxide particles became a popular method as it offered control over morphology and composition depending on the precursor, temperature and density of the fluid. Under supercritical conditions, reaction kinetics is very high and the formation of metal oxide particles occurs rapidly due to its low solubility. In previous studies, Fe(NO3)3, Fe2(SO4)3, FeCl2 and Fe(NH4)∙2H∙(C6H5O7)2 were used as precursors in forming various iron oxides in supercritical water. Fe(NH4)2H(C6H5O7)2 formed spherical magnetite (Fe3O), while all others formed spherical hematite (α-Fe2O3).  Both in the sol-gel method and rapid hydrothermal process, the metal oxide goes through two reactions; first the metal salt undergoes hydrolysis in which the salt will react with water to form a hydroxide and second, a dehydration process where the hydrogen is separated to form an oxide. Dehydration occurs from the outer surface of the hydroxide particle and therefore smaller particles may favor faster dehydration rates (Adschiri et al., 1992). These reactions are given in equations 4-1and 4-2 respectively.  4-1  4-2  In equation 4-1, B is the anion of the ferrous salt which is either Cl- or SO4- in this study. Due to excess hydrogen ions, the solution is slightly acidic and the salt is said to be a lewis acid.  47  At higher temperatures, dehydration occurs in which the iron is oxidized from Fe2+ to Fe3+. This forms magnetite if the environment is anaerobic while forming hematite or other hydroxides in an aerobic environment. This is known as Schikorr reaction and is a common reaction which takes place in boiler tubes forming magnetite (Shreir et al., 1994). 4-3  A residence time of approximately 2 minutes yielded 50 nm diameter hematite particles from Fe(NO3)3 in supercritical water. Adshiri (1992) also tested reaction rates and found that it formed hematite nano-particles even when the reactor residence time was cut down to less than 1 second although particles were smaller (<20 nm) suggesting that reaction times are very high (Adschiri et al., 1992). Synthesis in supercritical conditions occur faster partly due to the lower dielectric constant of water and the resulting enhancement of the reaction rate according to Born’s theory (Adschiri et al., 2000).  Although Aschiri had used ferrous chloride in his experiment, it is believed that the system still contained oxygen, making it a highly oxidizing environment under supercritical conditions. This would explain why hematite was formed in Aschiri’s experiment with ferrous chloride.  4.4 Experimental Procedure 4.4.1 Experiments with Precursor Before each experiment, the primary pump was turned on and the system was flushed with deionized water for about one hour. Both the primary and injection tanks were continually 48  sparged with nitrogen throughout the experiment to deaerate the system. The dissolved oxygen sensor was placed in the injection tank to deaerate the solution until it reached 0.00ppm which usually took one to two hours. Once the dissolved oxygen content reached the desired level in the injection tank, the DO sensor was moved to monitor dissolved oxygen levels in the primary tank for the remaining part of the experiment.  The back-pressure regulator was adjusted until the pressure at the transducer was reading the required pressure for the particular experiment. Although the pressure transducer is located before the heaters and several feet away from the test section, it is assumed to that the readings are representative of the pressure in the test section as the pressure drop between the two points is small and negligible compared to the system pressure of 24 MPa. Flow rates of the primary pump were measured and adjusted according to the required 3.0L/hr. The injection pump connected to a tank of deionized water was turned on and the new flow rate was adjusted to 3.5L/hr using the injection pump. The pressure remained constant during this time and the injection pump flow rate was calculated as the difference between the two flow rates. Once the desired flow rate and pressure were achieved, the heaters were turned on and allowed to heat the system to supercritical conditions. Temperatures of the bulk fluid inlet and outlet of the test section were checked until the system reached steady state temperature. The injection pump was momentarily turned off in order to switch the injection tank from deionized water to precursor solution. This point would be observed in the temperature measurements by a momentary increase in the temperature profile of the test section due to reduced flow rates. Once the tank with precursor solution was connected, the injection pump was turned on once again and the temperature of the test section dropped to its original temperature. The flow rate at the exit was  49  measured once more to check if they were consistent from before. Each of the experiments ran for approximately 40 minutes, with the exception of the low concentration ferrous chloride experiment which lasted for 400 minutes at one tenth the concentration. Therefore, the total mass of the salt injected was kept constant in all experiments.  For low heat flux, ferrous chloride experiment, a 1.0m, 1/8” x 0.028” outer diameter stainless steel 316L tube at 6V was used rather than the 1.8m at 12V used for the rest of the experiments. It was realized during this experiment that there was significant electrical resistance in the wiring and heat loss through the mounting blocks such that the temperature rise in the test section was very small. Therefore, for all subsequent experiments, the test section was modified to run at 1.8m and 12V yet maintain a maximum current under design ratings of 200A.  4.4.2 Experiments with Magnetite Particles Experiments with magnetite particle injection was also initially tested using commercially available magnetite powder by Sigma Aldrich (Iron (II,III) oxide, > 98% Fe3O4). The use of prepared magnetite powder for the experiments offered several advantages over using a precursor; the amount of magnetite particles in the system could be accurately known without the need to determine yield rates and the size of the particles could also be externally controlled. The magnetite powder had a diameter of 50 nm or less and was suspended in deaerated solution in the injection tank. The solution was continually mixed with an electrically powered mixer to prevent particle settlement. The magnetite particles were observed to agglomerate and stick to the glass flask and tubing, lowering the overall concentration of magnetite powder in the solution. To mitigate this problem, the pH of the solution was increased to 10 using sodium  50  hydroxide where the zeta potential of the solution/magnetite favored dispersion. However, the HPLC injection pump continued to experience difficulty pumping at consistent concentration of magnetite nano-particles, therefore further experiment with magnetite slurry was not pursued.  4.4.3 Post Experiment Cleaning After each experiment with the exception of experiments 7 and 8, the test section was removed and replaced with a temporary tube and the system was chemically cleaned. For experiments 7 and 8, both the test section tube and the rest of the system underwent chemical cleaning. The system was cleaned using 1N hydrochloric acid at 80⁰C for one hour before each test to ensure that magnetite deposits were removed from the system. After cleaning with hydrochloric acid, the entire system was flushed out with de-ionized water for several hours. When required, mechanical cleaning of tubes and fittings was also conducted by disassembling the sections and cleaning the section with an ultrasonic cleaner and/or a mechanical brush.  4.5 Analytical Methods 4.5.1 Online Thermal Resistance Monitoring The electrical heater was designed and constructed to serve several purposes. First, the heaters increase the temperature from subcritical to supercritical temperatures, allowing for a wide range of temperatures and heat flux conditions which can be tested. Second, it offered an online monitoring of fouling in the system. Deposits on the heat transfer surface would increase thermal resistance which would be detected by an increase in temperature of the tube outer wall. By modeling magnetite deposits and heat transfer characteristics, the deposition and removal rates  51  can be calculated. Calculated values can be compared to the deposition found in SEM images and cleaning methods.  The test section is electrically wired with current running directly through the tube causing the electrical resistance of the 1/8” tube to dissipate power in the form of heat. It is assumed that the electrical resistivity remains constant along the tube and therefore the heat flux into the water will also remain constant along the tube. This method for heating the tube was chosen over heating tapes and ceramic heaters for its uniform heat flux demonstrated by Teshima (1997) and Khan (2005). The test section tube is connected to a 2400VA transformer with its primary coil wired in series and secondary coil wired in parallel using 6V center taps. Power input was provided by two of three legs of a three-phase delta-configured power supply, which supply a voltage of 208V RMS to the transformer. Fuzzy Pro logic controllers monitored the temperatures for both the preheaters and test section. These have PID control which read temperatures from the thermocouples and send a signal for optimal power level to the solid-state rectifiers (SSR) which act as a gate for the electrical current.  Metal oxide varistors (MOVs) were installed both at the inlet before and after the SSR to prevent voltage spikes from damaging the SSR and transformer. The transformer connects to the tube with one connection to the center and two ground wires to each of the ends. This is to prevent any ground loops from forming within the system, which may create safety hazards. The resistance of the 1/8” SS316L tube was measured to be on average ~0.13 Ω per meter, however the effective resistance is reduced to a quarter of its value due to the parallel connection.  52  Thermal expansion of the tubing was also considered which can be as high as a few millimeters from room temperature to experimental temperatures. The heat exchanger is free to move axially a short distance, allowing thermal expansion to occur without introducing significant stress on the test section.  For both safety and protection of the equipment from voltage on the tube, secondary grounding wires were attached at the inlet of the injection line and after the heat exchanger. Three tube mounting blocks were constructed from SS303 for clamping the tube and the electrical conductor together. These mounting blocks were specially designed and optimized to offer high thermal resistance to decrease heat loss from the test section yet maintain high electrical conductivity for the electrical path. Aluminum heat sinks were attached to the mounting blocks to dissipate any excess thermal heat. The configuration of the test section and mounting blocks are provided in Figure 4-3.  53  From main pump  Test section From injection pump  Wall temperature thermocouple  Mounting blocks Tee with inlet bulk thermocouple  Heat exchanger  Figure 4-3: Mounting block and test section inlet  Temperature along the tube was measured using K-type Chromel/Alumel thermocouples with twisted shield for protection against electrical noise. Although not as accurate as RTD, they are very robust even at high temperature and can be calibrated to be accurate within less than 1⁰C. Six thermocouples were connected to the control panel for input power control and safety reasons. Eleven thermocouples were spot welded to the 1/8” test section tubing on the outside wall. These thermocouples were placed on the top of the tube, however, since buoyancy is negligible in these conditions, the location around the circumference of the tubing should not have a significant impact on the temperature reading. Thermocouples were clamped onto the  54  outside wall of the outlet of heaters 1 and 2 and the tube downstream of the heat exchanger of the main line. Bulk fluid temperatures were measured using Inconel sheathed, ungrounded probes which were inserted at the Tees with the probe end approximately at the center of fluid flow. The outlet bulk temperatures were fed into the control panel and were used to control the temperature of the electrically heated section. A wiring and thermocouple placement diagram is shown in Figure 4-4.  AC voltage and current into the heated tube sections were measured for calculating power input. Although these measurements are not used for determining the enthalpy of the water, it provided a secondary confirmation for ensuring temperature rise corresponding to the electrical power provided by the system. A fluke clamp meter was used to measure AC current with resolution of ±0.1A. Measurements of the voltage across each half-section were performed with a Mastercraft multimeter with a resolution of ±0.01V.  Input power calculations derived from current/voltage measurements closely matched input power calculations using bulk temperature measurements, shown in the table below as the electrical and thermal power respectively. The small discrepancy between the two values is likely due to thermal loss to the mounting blocks and surrounding insulation, hence the electrical power is slightly higher than the actual heat gained by the fluid.  Table 4-1: Input power comparison Experiment 1  First half of tube  Second half of tube  Current [A]  Voltage [V]  Current [A]  Voltage [V]  Power [W] (electrical)  54.3  10.00  51.5  9.93  1054  Power [W] (thermal) 1025  55  Calibration of the thermocouples was conducted by measuring the outside wall temperatures with the bulk inlet and outlet temperature of the test section set at a saturation temperature and pressure, and the section was thermally insulated with no input power. This process was completed before the experiment each time a new test section was installed. Temperature was recorded for thermocouples and any significant temperature differences were noted. For the thermocouples, the temperature measured usually fell within one or two degrees of expected wall temperature. These were used to correct the temperature readings for any future measurements for each experiment.  Subsequently, the accuracy of the temperature profile calculated by the MATLAB simulation fluidproperties.m was checked by comparing the simulation to experimental temperature profile with inlet and outlet temperatures at 350⁰C and 400⁰C respectively. The measured and calculated temperatures using correlations of Swenson and Bazargan are shown in Figure 4-6.  56  Omega 1200 High Speed Isolated Measurement System  Flow  15 cm  Mounting blocks  Parallel  Transformer  Series SSR Fuzzy Pro logic controller  Fuzzy Pro logic controller  Figure 4-4: Electrical wiring, heater controls and location of surface welded thermocouples and bulk fluid thermocouples  57  Figure 4-5: Saturation temperature of 231.51ºC at 417psi, calibration for Experiment #1  Figure 4-6: Temperature of test section, calculated vs. calibrated thermocouples  58  It was observed from the comparison that Swenson’s correlation predicted temperatures well both in subcritical and supercritical conditions. Bazargan’s correlation over-predicted heat transfer enhancement at bulk fluid temperatures above 380°C. Therefore, all subsequent simulations presented in this study use Swenson’s correlation for determining the heat transfer coefficient. The temperatures read were quite stable (std = 0.7⁰C for experiment #1) and therefore, any temperature differences due to fouling can be measured with an accuracy to 1⁰C.  Temperature measurements of the test section thermocouples were recorded using a data acquisition hardware and computer. The Omega Multiscan 1200 has 24 electrically isolated channels and two internal channels which are used for cold junction compensation. The 16-bit temperature measurements offer 3.12 uV or 0.1⁰C resolution. Electrically isolated channels were needed for the temperature measurements since the thermocouples were welded and electrically connected to parts of the test section which are at various voltage potentials. AC line rejection and averaging helped to reduce noise caused by the AC voltage applied to the test section in the measurement. Temperature data is sampled at 20kHz frequency and the data was sent to “TempView” software, displaying real-time temperature measurements. The measurements were automatically averaged over a 5 second period and then outputted to the text file which was later analyzed in MATLAB.  4.5.2 SEM Imaging for Filters and Tube Deposits A Hitachi S-3000N in the Frank Forward Materials Engineering building at UBC was used for SEM imaging, with the exception of those photographs which are labeled as Hitachi 2300 SEM, also located in the Frank Forward building.  59  To observe the morphology of the deposit on the tube inner-wall surface, the test section was cut into 1.5cm long samples using a 1/8” tube cutter. The sample was then carefully cut in the middle of the tube by grinding halfway through the tube until there was exposure to the other inner wall. Then the tube was bent slightly to open the section and the sample was placed into a mounting clamp for SEM imaging.  Figure 4-7: Test section sample for SEM analysis  4.5.3 Particulate Filtering System Three filters were constructed and installed at various locations in the system to collect particles formed in supercritical water. After each experiment, the filters were removed and analyzed with SEM/EDX with the objective of determining particle size, shape, and composition.  The main body for the low temperature, high pressure (LTHP) filter holder shown in Figure 4-8 was constructed from stainless steel 316L. The filter was designed and constructed with the  60  intention of allowing future filter holders to be constructed from any material and allow for easy assembly and reassembly after each experiment (see Appendix B for details). The metal spiral wound sealant allows the filter to reach high temperature and pressure applications and compensates for uneven or scratched surfaces of the mating parts. Glass fiber filters were inserted into the LTHP filter to collect particles in the primary line after the heat exchanger as well as to protect the back pressure regulator which was sensitive to particulate fouling. After each experiment, the glass filter was carefully removed, left to dry overnight, and was gold coated using a gold sputtering machine for SEM/EDX analysis.  Figure 4-8: SS316L LTHP filter installed after the heat exchanger  61  A high temperature, high pressure (HTHP) holder was made from Swagelok 316L fittings shown in Figure 4-9. A stainless steel 304 perforated metal with a porosity of 22% and a hole diameter of 0.02” was fitted into the holder to provide back support for the filter. Glass fiber filters were initially tested in supercritical water, however, the filters were destroyed when exposed to supercritical water. It is thought that at high temperature and pressure, the glass fiber mechanical properties degrade significantly and the pressure drop across the filter may have caused it to break down. It was decided that the glass fiber filters would be used to collect particles only at low temperatures. In replacement of glass fiber filters, sintered silver membranes from SPI supplies with absolute retention pore size of 0.2 µm and thickness of 50 µm were used for high temperature applications. The thin filter makes it ideal for higher flow rates and the conductive properties of the silver membrane make it excellent for SEM/EDX and XRD analysis. The filter holder and its components are shown in Figure 4-9.  Figure 4-9: HTHP filter with silver membrane and SS304 back support  62  Finally, a polypropylene 25mm Swin-Lok filter holder was used for the low temperature, low pressure (LTLP) filter shown in Figure 4-10. The same glass fiber filters used in the LTHP filters were also used for the LTLP filter. A summary of the properties of the glass and sintered silver filters are described in Table 4-2.  Figure 4-10: Glass fiber membrane and Swin-Lok LTLP filter holder  Table 4-2: Glass and silver membrane filter comparison  Sintered Silver Glass Fiber  Commercial Supplier SPI Supplies Millipore  Effective Pore size [µm] 0.2 0.7  Thickness [µm] 50 380  Water Flow [mL/min/cm2] 17 1.4  Temp Max 550 500  SEM and EDX analysis were taken of clean filters to determine its composition, and therefore providing information about the background of the filters. This helps to identify which elements  63  found on the filter are coming from the particle and which are the background of the filter. The silver membrane was found to be composed of only silver and oxygen. The glass fiber filter had Si, O, Ca, K, Ti and Zn.  64  Figure 4-11: SEM photograph of clean silver membrane, 0.2µm pore size, using a Hitachi S3000N SEM  Figure 4-12: EDX of silver membrane to determine elemental composition of background  65  Figure 4-13: SEM photograph of clean glass fiber filter, 0.7µm pore size, a Hitachi S2300 SEM  Figure 4-14: EDX of glass fiber filter to determine elemental composition of background  66  4.5.4 Deposit Thickness and Strength of Oxide Adhesion To quantify the mass of magnetite deposited on the tube, two techniques for magnetite removal was used; mechanically using an ultrasonic cleaner and chemically using acid wash procedures. 5.0 cm sample tubes were cut at 30.0cm intervals along the test section for magnetite deposition quantification. For ultrasonic cleaning, the sample tube was immersed in 25.0mL of 1N hydrochloric acid at room temperature and cleaned in a Cole-Parmer ultrasonic cleaner for 20 minutes. The sample tube was then removed and the solution was heated on a bunsen burner with a watch plate and evaporated in a fume hood. Once the beaker was completely dry, it was washed with 25.0mL of 1% nitric acid solution. 2.0mL was taken from the sample and diluted in another sample container of 25.0mL, 1% nitric acid. This was repeated once more to achieve three sets of sample solutions which have of 1×, 13.5× and 182.25× dilution ratios.  For acid wash, the sample tube was immersed in 15.0mL of 1N preheated hydrochloric solution (≈75°C) for 2 minutes. The sample tube is then removed from the solution and the solution is further heated to evaporate it completely. The beaker was then washed using the same procedure as the ultrasonic cleaning method.  A Varian AA240 Atomic Absorption Spectroscopy (AAS) was used to analyze iron concentrations in the samples. AAS uses measured absorbance to determine analyte concentrations according to the Beer-Lambert law. Calibration of the AAS was conducted by first measuring the absorbance using standards of known concentration. For these experiments, the calibration was done with 0.0, 1.0, 2.0, 5.0, and 7.5mg/L of iron standards in 1% nitric acid  67  matrix. All sample solutions were also digested in a 1% nitric acid matrix. The detection limit of atomic absorption spectroscopy for iron is approximately 0.1 mg/L (Willard et al., 1969).  Since stainless steel 316L also contains iron, there would be dissolution of iron from the tubing itself and not just the deposited magnetite. To account for this, clean tubes without deposits were exposed to the same conditions as those of the sample tubes and its iron content was measured using AAS. This was done for both the ultrasonic cleaning and acid wash procedures and the resulting iron concentration from the clean samples were subtracted from the raw sample tube concentration.  A mass balance of the entire system was pursued by measuring the concentration of ferrous species in the effluent, filters and test section. Effluent concentrations were determined by taking 20.0mL samples and analyzing them using Atomic Absorption Spectroscopy in a similar manner to the test section; first by dissolving any possible magnetite that might have passed through the filter, evaporating the solution, and re-dissolving in 1% nitric acid matrix. For experiments with ferrous chloride, the effluent samples were taken during the experiment directly from the system effluent. It was determined that due to changing deposition/exfoliation rates over time, an averaged effluent concentration would be more representative for mass balance purposes. Therefore for all ferrous sulfate experiments, the effluent was stored in the high-densitypolyethylene (HDPE) tank until the experiment was completed and a sample was taken of the total effluent. In addition, for all ferrous sulfate experiments, sulfate was known to interfere with the AAS in measuring iron, therefore as a way of mitigating this, a copper sulfate of 0.1 mol/L and 1% nitric acid matrix was used for these standards and sample solutions.  68  5 Results and Discussion of Experiments A total of eight experiments and one blank were conducted using the SCW once-through flow apparatus. After each experiment, filters, effluent, and tube samples were analyzed using SEM and AAS to determine significant findings. A brief summary of the results is provided in Table 5-1 while a more detailed overview and collection of all results can be found in Appendix C.  The two precursors used were ferrous chloride tetrahydrate (FeCl2·4H2O) from Fisher Scientific and ferrous sulfate heptahydrate (FeSO4·7H2O) from BDH. The experiments are broken down into the following categories: reference condition, effect of heat flux, high pH, low concentration and blank run with no precursor.  Table 5-1: Summary of experiments  Run  Category  Pressure [MPa]  Tube length [m]  Bulk Temperature Inlet [⁰C]  Outlet [⁰C]  Heat Flux [kW/m2]  pH  Concentration [mmol Fe2+/L] Dissolved Fe2+  Precursor  1  Reference  23.7  1.8  350  400  102  3  5.48  FeCl2·4H2O  2  Reference  23.9  1.8  350  397  97  3.5  4.97  FeSO4·7H2O  3  Heat Flux  22.4  1.0  372  376  24  3  5.31  FeCl2·4H2O  4  Heat Flux  23.0  1.8  371  376  13  3.5  5.04  FeSO4·7H2O  5  Heat Flux  23.3  1.8  384  383  -  3.5  5.32  FeSO4·7H2O  6  High pH  23.8  1.8  350  396  97  9  5.11  FeSO4·7H2O  7  High pH  23.7  1.8  200  370  92  9  5.17  FeSO4·7H2O  8  Concentration  23.7  1.8  350  395  96  3  0.56  FeCl2·4H2O  -  Blank  23.7  1.8  351  399  99  Neutral  0.00  -  69  5.1 Reference Condition Experiments were run with ferrous chloride and ferrous sulfate precursors with the test section at maximum heat flux of approximately 100 kW/m2. These were considered to be reference experiments that all other experiments could be compared to. Temperatures of the bulk fluid inlet and outlet were 350°C and 400°C respectively at 23.7MPa. The injection tank consisted of 1L of precursor in deaerated, deionized water while the primary tank had deaerated, deionized water, producing an acidic solution (pH≈3) once both fluids were mixed. The injection method is similar to those described by Adschiri et al. (1992) in hydrothermal synthesis.  5.1.1 Temperature Analysis The transient temperature of the test section outer wall for reference condition with ferrous chloride precursor is shown in Figure 5-1. As mentioned in Chapter 4, the large spike and subsequent fall in the temperature at 15-20 minutes is due to the momentary shutoff/on of the injection pump for switching the supply line from deionized water to ferrous chloride solution. At subcritical temperatures, the temperature remain constant throughout the experiment. At supercritical temperature (1.65m from the inlet), the wall temperatures increase rapidly at the start and reaches an asymptotic temperature after 10 minutes of injection.  70  Injection  Figure 5-1: Temperature measurements of test section vs. time for reference condition with ferrous chloride precursor (Experiment #1)  Fouling was observed by the sharp rise in temperature as well as an increase in system pressure to 26.2 MPa during the injection period as a result of restricted flow. The temperature profile at the 1.65m location has an asymptotic fouling behavior similar to those found in several studies of iron oxide fouling (Müller-Steinhagen et al., 1988; Newson et al., 1983; Thomas & Grigull, 1974). It was speculated that the removal rate increases as the deposition thickness increases, since the bond at the top layer of the surface weakens as the deposit grows higher and is more likely to be removed by the shear force of the fluid (Newson et al., 1983). Another explanation presented by Thomas and Grigull (1974) suggested that the initial high deposition rate was due to the particles adhering to the troughs in an initially rough tube. As deposition continued, the  71  troughs would be filled and the tube surface would become effectively smooth and more difficult for the particles to adhere to.  High fouling rates observed at the 1.65m location from the inlet was converted to an equivalent magnetite deposition thickness at the asymptotic temperature. This was found to be 0.47 mm or 1.5g of magnetite in a 15 cm long section of the tube. Such high fouling rates are unlikely with only magnetite fouling, and therefore the high increase in temperature may be a product of mixed recrystallization of ferrous chloride as well as magnetite particles. Such a scenario should be considered as the solubility of any salt, including ferrous chloride would be very low in supercritical conditions.  Temperature measurements for ferrous sulfate precursor showed variations in the test section temperature but also in the bulk fluid temperatures (Appendix C. 2). This may be due to inconsistent flow rate of the injection pump for this experiment. Once this was taken into consideration, it was concluded that no significant fouling was observed from the temperature measurements.  5.1.2 Filter Analysis In the experiment with a ferrous chloride precursor, it was determined after the experiment that the secondary line was at a significantly lower flow rate than initially set. The HTHP and LTLP filter was visibly unchanged and showed very few particles collected through SEM images. On the other hand, the HTHP filter for the ferrous sulfate precursor had a very fine but visible black film on the silver membrane after the experiment. SEM analysis of the HTHP filter revealed  72  significant amount of particles collected, with particle diameters ranging from less than a micron to several microns in diameter. These particles are much larger than those seen by Adschiri et al. (2000) in hydrothermal synthesis of iron oxide particles which observed the formation of ~50nm diameter particles. However, there is a possibility that the larger particles were captured by their system filter which had not been analyzed.  5.1.3 Deposit Analysis SEM images of the test section revealed that the ferrous chloride precursor produced magnetite particles early in the test section. The particles which deposited near the inlet were smaller with few larger micron-sized particles. Further down the test section at higher temperatures, the particles deposited grew in size to several microns in which deposition was dense and compact.  Figure 5-2: SEM photograph of magnetite deposit on test section tube surface, 0.15 m location. Reference condition with ferrous chloride precursor (Experiment #1)  73  Figure 5-3: SEM photograph of magnetite deposit on test section tube surface, 0.60 m location. Reference condition with ferrous chloride precursor (Experiment #1)  Figure 5-4: SEM photograph of magnetite deposit on test section tube surface, 1.20 m location. Reference condition with ferrous chloride precursor (Experiment #1)  74  Figure 5-5: SEM photograph of magnetite deposit on test section tube surface, 1.65 m location. Reference condition with ferrous chloride precursor (Experiment #1)  X-ray diffraction (XRD) was conducted on the oxide deposits found in the test section tube by mechanically removing the deposit. Figure 5-6 and Figure 5-7 show XRD results for ferrous chloride and ferrous sulfate precursors respectively. For ferrous chloride, a General Area Detector Diffraction System (GADDS) using a copper radiation operated at 40kV and 40mA scanned the deposit from 21.4⁰ to 71.8⁰ at 0.050⁰ increments. A diffracted beam graphite monochromator and Hi-Start detector was used for producing and detecting the signal. Signals were relatively weak and the peaks were broad due to the small quantity of deposit that was available from the tube (see Figure 5-6).  75  Figure 5-6: XRD of tube deposits of test section, reference condition ferrous chloride precursor (Experment #1)  Figure 5-7: XRD of tube deposit of test section, reference condition ferrous sulfate precursor (Experiment #2)  76  For ferrous sulfate reference experiment, a D-8 Advance X-ray Diffractometer using a copper anode tube operated at 40kV and 40mA, scanned the filter from 5.0⁰ to 70.0⁰ at 0.040⁰ increments. A diffracted beam graphite monochromator and NaI scintillation detector was used for producing and detecting the signal. XRD of the deposit indicated a mixture of magnetite and hematite which may have formed due to oxygen contamination during the experiment (see Figure 5-7).  To determine the deposition thickness along the test section, cleaning methods described in section 4.5.4 were applied to the sample tubes. This is similar to chemical analysis methods for deposition determination used in previous studies by dissolving magnetite layers from experimental tubes using hydrochloric acid (Newson et al., 1983).  Between ferrous chloride and ferrous sulfate precursors, the amount and location of the deposit was relatively similar. At lower temperatures near the inlet, the deposit was found to be less and the deposit thickness grew as the temperature increased towards the outlet of the test section. Ultrasonic cleaning was able to remove the majority of the deposit, however, some of the deposit remained intact to the tube and was only removed after the acid wash procedure. It was observed that for tube sections which were subject to supercritical temperatures, depositions remained intact and were much more difficult to remove than those exposed to subcritical temperatures. This suggests an increase in the oxide adhesive strength to the surface at supercritical conditions, shown by the increased amount of magnetite which was removed only after an “aggressive” cleaning method, regardless of the precursor injected.  77  In a study conducted by Yeon et al. (2006), the deposition characteristic of hematite particles onto a heated zircaloy tubing was found to change by adding ferrous species to the solution. Under single phase flow conditions, there was almost no deposition of hematite particles onto a zircaloy surface unless the particles were mixed with ferrous ions. The deposit was found to form two layers on the surface; a top layer which was easily removed with ultrasonic cleaning using distilled water and an inner layer which was only removed after ultrasonic cleaning with concentrated acid. Yeon et al. (2006) concluded that the top layer was formed from particle deposition while the inner layer was formed from the precipitation of ferrous ions. Therefore, the crystallization of dissolved ferrous species formed a stronger bond to the surface which then allowed the loosely attached particles to stick to the wall. In the case of the ferrous precursors in supercritical water, the higher strength of the oxide in supercritical conditions is believed to be due to small amounts of crystallization occurring at the interface between particles and the tube surface due to the solubility drop of magnetite in supercritical water.  78  Figure 5-8: Deposit thickness on test section for reference condition with ferrous chloride precursor (Experiment #1)  Figure 5-9: Deposit thickness on test section for reference condition with ferrous sulfate precursor (Experiment #2)  79  5.2 Effect of Heat Flux Experiments were run with low heat flux and no heat flux conditions and are compared with the reference condition. Low heat flux of approximately 20 kW/m2 in the test section was conducted for both ferrous chloride and ferrous sulfate precursor, with temperature ranging from 371°C to 376°C. A third experiment under no heat flux condition had ferrous sulfate precursor solution and the temperature was at 384°C.  5.2.1 Temperature Analysis Temperature measurements were taken for the low heat flux conditions, however, no significant differences in temperature were found in these cases.  5.2.2 Filter Analysis SEM analysis showed sub-micron to several micron-sized particles scattered along the membrane similar to those seen in the reference conditions for ferrous sulfate (see Figure 5-10). The filter shows particles of various shapes and sizes, some which appear to be crystalline. 20.0keV area EDX was applied on the filter which confirmed the presence of iron oxide along with silver from the background as shown in Figure 5-11. The absence of other metal oxides such as nickel, chromium, or molybdenum oxides support the assumption that corrosion of the stainless steel 316L test section under experimental conditions is negligible. The HTHP filter for ferrous sulfate at low heat flux was fairly uniform magnetite particles of 1 µm size with agglomerates of smaller particles on HTHP filter suggesting that the type of precursor has some influence on the particles formed under the same conditions.  80  Figure 5-10: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for low heat flux with ferrous chloride precursor (Experiment #3)  Figure 5-11: EDX of HTHP filter for low heat flux with ferrous chloride precursor  81  Figure 5-12: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for low heat flux with ferrous sulfate precursor (Experiment #4)  Figure 5-13: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for no heat flux with ferrous chloride (Experiment #5)  82  For the experiment run with ferrous sulfate with no heat flux and supercritical temperature, the particles on the HTHP filter were of larger diameter than in low heat flux experiments and some of these larger particles also appear to have an octahedral structure as seen in Figure 5-13, which is a common crystal structure for magnetite (Anthony et al., 1997). XRD was carried out on the HTHP silver membrane filter of Experiment #3 to determine the crystal structure of the deposited particles. Results show that the peaks match very well with those of magnetite and comparison of these peaks to other iron oxide such as hematite indicated that these were not present in the sample. Carbon graphite is also present in the sample and is believed to have been detected due to the carbon tape used on the filter for SEM.  Figure 5-14: XRD of deposit on HTHP filter for low heat flux with ferrous chloride precursor (Experiment #3)  83  5.2.3 Deposit Analysis Deposition thickness of the test section for ferrous sulfate precursor under low and no heat flux conditions were examined and the results are found in Appendix C. SEM images of the tubes show very little deposition and these findings are confirmed using cleaning methods.  Figure 5-15: Deposit thickness on test section, low heat flux with ferrous sulfate precursor (Experiment #4)  The effect of heat flux compared to reference condition suggests that lower heat flux results in lower deposition. This compares to previous models which showed that deposition had a direct influence from heat flux (Electric Power Research Institute Inc., 2003). In previous studies, experiments showed that deposition of magnetite particles in bulk concentration of 0.16mg to 5.04mg of Fe kg/L was 30% lower on an unheated surface than one with heat flux of 400kW/m2 (Newson et al., 1983).  84  Although the original Kern and Seaton model does not account for the influence of heat flux on particle deposition, there have been several attempts in later studies to include heat flux by incorporating thermophoresis and an Arrhenius term with an activation temperature (MüllerSteinhagen et al., 1988). Thermophoresis has been shown to create a measurable change in the deposition of a particle onto a heated surface for particles as large as 11 µm (Epstein, 1997), although it primarily observed to influence smaller submicron particles. The thermophoretic velocity  is dependent on the fluid properties and temperature gradient and can be calculated  using the equation below:  5-1  where  is a coefficient determined by:  5-2  The constant k is 1.8 for gases and 0.26 for liquids, λp and λl represent the conductivity of the particle and liquid respectively (Epstein, 1997). The overall deposition velocity accounting for thermophoresis is described in equation 5-3:  (  )  5-3  The equation resulted in values for the thermophoretic velocity of Kth in the range of 10-10 m/s, much smaller than the mass transport velocity of 10-4 m/s for these experiments, suggesting that thermophoresis should have an insignificant impact on deposition.  85  In addition to thermophoresis, high heat flux conditions usually result in higher surface temperature given constant flow rates, which provide the activation energy required for the particle to stick to the wall. In an experiment with alumina particles depositing on a heated surface, Müller-Steinhagen et al. (1988) observed that an asymptotic fouling resistance had a maximum at any given heat flux for a fixed flow rate, concentration and bulk temperature. To model the data, Müller-Steinhagen included both thermophoresis and activation energy into the Kern and Seaton model. The model showed that the increase in heat flux initially resulted in the increase of surface temperature resulting in higher deposition rates. However, as Tw increasingly exceeds Tb, deposition is counteracted and eventually overtaken by thermophoresis thereby decreasing deposition with increasing heat flux. In general, the correlation over-predicted deposition velocities and a correction factor had to be introduced to give a reasonable fit to the experimental data. Such a procedure demonstrates the difficulty in accurately modeling the complex nature of particles in heat transfer flow.  Although the Arrhenius term predicts lower deposition rates for lower heat flux conditions, it fails to adequately explain the extent to which deposition rate decreased in these experiments compared to reference conditions. Müller-Steinhagen et al. (1988) postulated that surface/fluid temperature may also affect the structure of the deposit, making it more porous and loosely packed in certain conditions. It is not known if this is affected by absolute wall temperature or by the wall-liquid temperature difference. If it is assumed that the lower heat flux results in loosely packed deposits which can easily be subject to removal, this may explain the lower deposition of the low heat flux condition observed in this study. However, further research would be required  86  in order to make any conclusions regarding if heat flux has a direct influence on the deposit structure.  5.3 High pH Conditions Current CANDU reactors normally operate with a pH of 9 ~ 10 using a simple water chemistry of lithium hydroxide for pH control (Burrill, 2000). Two experiments were conducted with ferrous sulfate as a precursor but at a higher pH of 9 (at room temperature). The experiments were run with temperatures running at 350°C to 400°C and 200°C to 370ºC.  The pH of the system was increased by adding sodium hydroxide salt to the feedwater. At first, ferrous sulfate and sodium hydroxide was mixed in the injection tank and the pH was measured directly. However, the introduction of sodium hydroxide to a ferrous sulfate solution immediately produced particles in the system which would cause problems with the injection pump. As an alternative approach, sodium hydroxide solution was introduced in the primary tank and allowed to mix with the ferrous sulfate in-situ of the system. Due to size limitations of the main tank, sodium hydroxide was added to the first tank after heating the entire system. The sodium hydroxide was weighed and mixed into a deaerated tank and a sample was drawn from the tank. The conductivity was measured and compared to the amount of sodium hydroxide required using a calibration curve shown in Figure 5-16. A linear regression gave the equation 53 for the concentration as a function of conductivity.  5-4  where C is the concentration [mmol/L] and σ is the conductivity [mS/cm].  87  Figure 5-16: Calibration of the conductivity meter for determining NaOH concentration in primary tank  The pH could not be measured directly in the system due to the high temperature and pressure of the solution at the mixing tee. Instead, the pH of the ferrous sulfate solution was determined by using data from Arden (1950) for ferrous sulfate and sodium hydroxide solution. Conditions in Arden’s experiment are similar to the conditions in this study. The solution was deaerated to eliminate the presence of oxygen and ferric ions which have been found to create large errors in the pH measurements. Oxygen in the system promotes the formation of ferrosic hydroxide, a ferric ion compound instead of the desired ferrous hydroxide (Arden, 1950).  The conductivity of the primary tank after dissolving sodium hydroxide was 2.02 and 2.10 mS/cm which correspond to a concentration of 10.49 and 10.91 mmol/L for experiments #6 and  88  #7 respectively. The ratio of Na to Fe was calculated with equation 5-4 taking into consideration the dilution of the primary stream with the injection stream. ̇  ̇ ̇  ̇  5-5  The molar ratio of Na to Fe was compared to data with Arden (1950) and the pH was determined to be approximately 9 in both experiments, as shown below.  Figure 5-17: Predicted pH of the system using data from Arden (1950)  5.3.1 Temperature Analysis For the experiment with the test section heated from subcritical to supercritical, large temperature changes were found at locations 0.15 m and 1.20 m while a steady temperature rise was found near the outlet at 1.65m. The temperature at the 1.20 m location approached 500°C  89  after 30 minutes which was considered too high and the power to the test section was momentarily turned off to cool the section. Once the temperature dropped to below 400°C, the test section was turned on once more to continue the injection at the original temperatures until the end of 40 minutes.  Injection  Figure 5-18: Temperature measurements of test section vs. time for high pH with ferrous sulfate precursor (Experiment #6). Power turned off briefly at t = 48 minutes.  Temperature measurements at 0.15 m from the inlet suggest that deposits exfoliated from the tube surface relatively easily and quickly. These deposition-removal cycles found in the temperature was also observed by Khan (2005) with the injection of sodium carbonate in supercritical water. These occurred with combined particle and crystallization deposition and signaled a weaker adhesion to the tube wall than pure crystalline deposit.  90  5.3.2 Filter Analysis Observing the particles captured on the HTHP and LTHP filters under an SEM, revealed that the particle sizes are consistently much smaller than in the reference case, falling in the 100-200nm diameter range. Formation of smaller particles in higher pH solution are confirmed by literature studies using sol-gels. Sugimoto and Matijevic (1980) displayed that under excess OH- ions, the magnetite particles formed were small, tens of nanometer in diameter and had a cubic morphology. On the other hand, large magnetite particles of 1.1 µm formed under 0.03M excess Fe2+ conditions.  The LTHP filter for both experiments had a significant amount of small particles deposited on the glass fiber. In experiment #7, it can be seen that the deposit was enough to completely cover all signs of the glass fiber matrix.  Figure 5-19: SEM photograph of particles collected on HTHP filter, 10.0k magnification, for high pH, ferrous sulfate (Experiment #6) 91  Figure 5-20: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH, ferrous sulfate (Experiment #6)  Figure 5-21: SEM photograph of particles collected on HTHP filter, 5.0k magnification, for high pH, subcritical, ferrous sulfate (Experiment #7)  92  Figure 5-22: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH, subcritical, ferrous sulfate (Experiment #7)  5.3.3 Deposit Analysis SEM images of the tube surface after the test section was disassembled confirm that there was very few particles attached to the surface. Furthermore, deposit thickness determination using cleaning methods indicate that the deposit is less than 1 µm shown by Figure 5-23.  93  Figure 5-23: Deposit thickness on test section (Experiment #6)  As mentioned in Section 2.3.3, the particles that deposit on the tube are submerged in the viscous sub-layer which dissipates turbulent bursts responsible for particle removal. The viscous sublayer thickness was calculated for the high pH at supercritical experiment using equation 5.5 below (Cengel & Cimbala, 2006).  5-6  The calculations showed that this distance was approximately 25 µm at 200°C, 14 µm at 350°C, and 6 µm at 400ºC under the experimental conditions. In the case of Experiment #6 at 0.15m location, the deposition-removal cycles had an average temperature change of 12°C from the initially clean state. This corresponded to a calculated thickness of the deposit right before the removal of 42 µm.  94  The cool down procedure was repeated in a similar manner for all experiments for direct comparison between the experiments. Significant amount of particles were found in the effluent of the system during the cooling/depressurization for both high pH experiments, indicating that removal rates post-injection were much higher for this deposit. This may be attributed to the increase in flow rate during depressurization. Turner et al. (1990) suggested that the change in flow rates are a larger factor in removal rates compared to steady state conditions.  5.4 Low Concentration and Blank Experiment The experiment with lower concentration of 0.5 mmol/L and proportionally longer duration of 400 minutes was conducted to determine the effects of concentration on deposition without changing the total amount of ferrous species injected. All three filters showed ferrous-chromium exfoliation occurring in the system indicating corrosion of the tube. Hence, the results from this experiment were not analyzed any further.  A blank test was also run with reference conditions but without a precursor. SEM analysis of the HTHP and LTLP filters showed almost no particles collected as expected. The LTHP filter had few particles which likely exfoliated from the heat exchanger, but the amount was insignificant compared to the experiments with a precursor.  5.5 Overall Summary and Mass Balance A summary of the significant results from each of the experiments are provided in Table 5-2.  95  Table 5-2: Results summary table - Part I  Parameter  Precursor  Filters HTHP  LTHP  LTLP  Tube Samples SEM  1  Reference case for ferrous chloride  FeCl2  Almost no particles  Large amount of submicron and some micron sized particles  Almost no particles  Small particles at inlet, large micron sized particles at high temperatures  2  Reference case for ferrous sulfate  FeSO4  Nanometer to several micron  Very Small particles of unknown size  Almost no particles found  Small and few deposits at inlet, large deposits at center of test section  3  Low heat flux  FeCl2  Several micron size  Very Small particles of unknown size  N/A  consistent micron sized particles found at inlet and outlet of test section  4  Low heat flux  FeSO4  Consistent ≈ 1 micron size  Very Small particles of unknown size  Almost no particles found  Very few deposits  5  No heat flux, all supercritical  FeSO4  Hundred nanometer to several micron size  Approximately 500nm consistent size  Almost no particles found  Combination of submicron and micron sized particles  6  High pH  FeSO4  Approximately 100-200nm consistent sizes  Approximately 100-200nm consistent sizes  Small amount of particles, hundrednanometer sized particles  Very few deposits  7  High pH, all subcritical  FeSO4  Approximately 100-200nm consistent sizes  Approximately 100-200nm consistent sizes  Almost no particles found  N/A  8  Low Concentration  FeCl2  Nanometer to several micron  Large amount of particles and flakes  Large amount of very small particles  N/A  Almost no particles  Few particles of sub-micron size  Almost no particles  N/A  Blank  -  96  Table 5-3: Results summary table - Part II Temperature Measurements  Deposit Thickness  1  Temperature rise at supercritical temperatures, no temperature difference for subcritical  10µm maximum  More depost strength in supercritical conditions  Recovery of 79%  Analyzed tube deposits, found magnetite  2  No difference found  12µm maximum  More deposit strength in supercritical conditions  Recovery of 82%  Analyzed tube deposits, found magnetite and some hematite  3  No difference found  N/A  N/A  Recovery of 52%  Analyzed HTHP Filter, found magnetite  4  No difference found  Less than 1µm thickness  No difference  Recovery of 72%  N/A  5  N/A  Less than 1µm thickness  No difference  Recovery of 77%  N/A  6  Temperature rise at supercritical temperatures, oxide deposition and exfoliation found at subcritical  Less than 1µm thickness  No difference  Recovery of 53%  N/A  Large amount of black particles found in the effluent after cooling the system  7  No difference found  N/A  N/A  Recovery of 39%  N/A  Did not dissassemble test section for this experiment  8  No difference found  N/A  N/A  Recovery of 83%  N/A  Did not dissassemble test section for this experiment  No difference found  N/A  N/A  N/A  N/A  Did not dissassemble test section for this experiment  Deposit Strength  Mass Balance  XRD  Comments Secondary line flow rate was much lower resulting in very few particles on secondary line filters  Test section is 1.0m long, test section not examined using cleaning method  97  A mass balance calculation was completed for all experiments, however, not all sections for each experiment were analyzed and therefore, are presented to provide an approximation only. The majority of the experiments had good recovery of the ferrous species injected into the system with the exception of experiment #7 which was below 50%. To determine the yield of magnetite for input into the simulation, the iron content found in the effluent was subtracted from the total injected amount. Hence, it was assumed that the iron mass which was not recovered had deposited as magnetite particles onto other surfaces within the system not analyzed, such as the heat exchanger or the tube fittings between the mixing tee and test section. Mass balance for future experiments may be improved by cleaning the heat exchanger after each experiment and analyzing the iron concentration in the cleaning solution. The equation for recovery and yield are given in equations 5-7 and 5-8 respectively.  5-7  5-8  In experiment #6 and #7, it was found that the effluent contained significant amount of oxide particles which passed through the LTHP filter, perhaps due to overloading. The experiment produced a higher yield of iron oxide particles indicated by the relative number of particles collected on the HTHP filter. To gain insight into the yield in the high pH experiments, the effluent from experiment #7 was first filtered using a Fisherbrand filter (Grade - P2) to separate the residue from the filtrate. The residue and filtrate were independently analyzed using AAS. It was found that the effluent was 57% particulate matter. Therefore, for experiments #6 and #7,  98  the same percentage of particulate matter in the effluent was assumed and the amount was added to the total yield of particles in the system.  Table 5-4: Mass balance of experiments #1 - 8 Experiment #  Fe [mmol/L]  Flow Rate [L/hr]  Run Time [min]  Effluent [mg]  Test Section [mg]  Total Filter Deposit [mg]  Total Iron Expected [mg]  Total Iron Measured [mg]  Recovery [%]  Estimated Yield  1  5.48  3.54  42  353  244  0.69  758  598  79%  52%  2  4.97  3.54  39  141  380  1.28  639  522  82%  78%  3  5.31  3.60  44  404  NM  NM  783  404  52%  48%  4  5.04  3.54  40  478  11.5  NM  664  478  72%  28%  5  5.32  3.50  40  536  26.1  NM  692  536  77%  23%  6  5.11  3.50  37  310  12.0  2.53  615  324  53%  78%  7  5.17  3.47  40  260  NM  NM  667  260  39%  83%  8  0.56  3.50  400  605  NM  NM  729  605  83%  17%  2+  99  6 Results and Discussion of Simulations 6.1 Comparison between Simulation and Experimental Simulations were run with parameters from experiments and these were compared to experimental deposition thickness obtained from the cleaning method. The magnetite concentration used for the simulation was obtained from the yield estimated from mass balance in Table 5-4, while the average particle size was estimated from the SEM images of the tube deposit and filters. For the reference conditions, the average particle size was estimated to be 4 µm and 1 µm for ferrous chloride and ferrous sulfate precursors respectively. For these reference condition experiments, the simulation results predicted magnetite deposition reasonably well. The simulation predicted similar deposition thickness for each case, however, with larger particles the peak deposit shifted closer towards the inlet of the test section.  (a)  (b)  Figure 6-1: Comparison between simulation and experimental for reference conditions (a) Experiment #1, ferrous chloride with average particle size = 4 µm, magnetite concentration = 200 mg/L (b) Experiment #2, ferrous sulfate with average particle size = 1µm, magnetite concentration = 299 mg/L. 100  For low heat flux conditions, the deposition thickness was overestimated by nearly a factor of four. However, the simulation did predict a lower deposition rate compared to the other experiments and a uniform deposition rate along the test section compared well to experimental results, as shown in Figure 6-2.  (a)  (b)  Figure 6-2: Comparison between simulation and experimental for ferrous sulfate experiments with (a) Experiment #4, low heat flux with average particle size = 1 µm, magnetite concentration = 108 mg/L and (b) Experiment #5, no heat flux with average particle size = 2 µm, magnetite concentration = 89 mg/L.  Finally, for the experiments at high pH, the simulation greatly overestimated the deposition thickness. This is likely due to the removal that occurred during and after the deposition process described in section 5.3.3. Higher removal rates in the high pH conditions suggest that the deposit had a weaker adhesion to the tube. As discussed in section 5.1.3, the strength of oxide adhesion was enhanced with the precipitation of ferrous species, and one explanation is that the 101  reduced amount of dissolved ferrous ions in the solution resulting from the high yield of particles formed under high pH condition may have caused the formation of a weak deposit. A second possibility is that the formation of molten sodium hydroxide salts had formed and deposited along with the magnetite particles in supercritical conditions. During cool down, the molten salt would have re-dissolved in the water, possibly carrying the magnetite particles along with it.  Figure 6-3: Comparison between simulation and experimental for Experiment #6, high pH, ferrous sulfate. Simulation was run with an average particle size = 0.15µm, magnetite concentration = 321 mg/L.  In all cases, the simulation overestimated the deposition of magnetite suggesting that surface attachment and/or removal rates cannot be neglected in magnetite particle deposition in supercritical water.  102  6.2 Simulation for CANDU SCWR As one of the design criteria in the proposed SCWR, the maximum fuel cladding temperature cannot exceed 850⁰C (Naidin et al., 2009). This is to ensure that the material’s mechanical and chemical properties can withstand the aggressive environment of supercritical water. Li et al. (2009) developed a 3-dimensional heat transfer simulation for a CANFLEX fuel bundle in a SCWR which showed that the maximum temperature of the cladding can reach as high as 802.2⁰C at certain locations of the fuel bundle.  The CANDU SCWR deposition simulation developed in this study does not account for oxide growth from the fuel cladding. This is due to the fact that the alloy for the fuel cladding has not been decided and the rate of oxide growth from the alloy varies dramatically depending on the material (Was et al., 2006). Additionally, only the surface temperature of the fuel cladding exposed to supercritical water is calculated since the design criteria of maximum fuel cladding is concerned with the surface temperature (Li et al., 2009).  Figure 6-4 shows the expected fuel cladding-surface temperature with a heat flux of 1000 kW/m2 with no deposition (see Table 3-1 for simulation parameters). The fluid enters the test section at 350ºC and 25MPa, saturated with 11µg/L of ferrous ions. As the temperature increases beyond the pseudocritical temperature, the solubility drops in accordance with Figure 2-3 and particles form in the bulk fluid. Figure 6-5 shows the temperature profile of the fuel bundle after one year of operation assuming a particle diameter of 150 nm.  103  Figure 6-4: Bulk fluid and fuel cladding-surface temperature without particle deposition  Figure 6-5: Bulk fluid and fuel cladding-surface temperature after 1 year of operation with magnetite particles of 150 nm diameter  104  Figure 6-6: Magnetite deposition thickness on fuel cladding after 1 year of operation for 150 nm and 1 µm average particle diameters  In Figure 6-6, the effects of particle diameter is compared between 150 nm and 1 µm. The experimental results from section 5.3 suggest that at a pH ≈ 9-10 where a typical CANDU reactor operates (Burrill, 2000), the formation of small, ~100 nm particles are likely. For the 1 µm scenario where 0.279 ≤ tp+ ≤ 8.284, inertial coasting is the dominant mechanism for deposition and the deposition velocity is approaching maximum rate. On the other hand, for the particle diameter of 150 nm where 0.006 ≤ tp+ ≤ 0.215, the deposition mechanism changes from diffusion to inertial coasting inside the reactor core and the deposition velocity remains closer to the minimum of the deposition velocity curve.  Maximum temperature difference between fouled and un-fouled cladding-surface temperatures was found to be 23.9⁰C occurring approximately 3 m downstream of the inlet.  105  Since the simulation was based from 1 fuel element, the cladding temperature represents an average among all 43 actual elements. The Li et al. (2009) simulation which looked at all 43 elements, found that cladding-surface temperatures can vary by several hundred degrees depending on the fuel element. The highest fuel cladding temperature of 802.2⁰C occurred at the 4 m location where an increase of 19.6⁰C due to fouling was found by this study. Therefore after several years of operation, fouling on the fuel cladding may cause temperatures to exceed the design criteria of 850⁰C if magnetite deposition considerations are not taken into account.  Calculation of the deposit thickness found in the simulation by Burrill (2000) discussed in section 2.3.4 suggest a peak deposit of 297 µm assuming mass transport of ferrous ions in a CANDU SCWR after one year of operation. This is in relatively good agreement to the results presented in this study considering several key differences in the parameters and assumptions in the model; a slightly higher saturation concentration of 18µg/L (Burrill, 2000) compared to 11 µg/L, temperature of core outlet of 384°C (Burrill, 2000) compared to 625°C, and the use of an ionic diffusion model.  Similarly, the transport and deposition model of Cook and Fatoux (2009) concluded a deposit thickness of 22.8 µm near the outlet of a CANDU SCWR reactor. The arbitrary deposition velocity assumed by Cook and Fatoux of 10-5 m/s which is slightly lower than the 4.2×10-5 – 28.7×10-5 m/s range for the deposition velocity calculated here using mass transport for 150 nm sized particles. Despite these differences, the maximum deposition thickness of 50 µm found from this study fits reasonably well with the model by Cook and Fatoux.  106  6.3 Limitations of the Simulation In Chapter 5, surface attachment resistance and removal rates were found to be important factors in the deposition of magnetite particles in supercritical water. However, these were neglected in the simulation as further experimental work is required to properly include a surface attachment and removal term. Therefore, for the situation where surface attachment and removal terms are unknown, mass transport models provide a relatively good estimate on the upper limit for deposition rate. Similar results were found by Turner and Klimas (2000) who examined magnetite particle deposition onto Alloy 600 under single-phase forced convection at high temperature and pressure water. Significant deviations were found between calculated and measured deposition velocities, all of which the mass transport model overestimated deposition rates. Surface attachment was determined to be the main limitation in this experiment while removal was found to be insignificant. Turner measured the isoelectric point for both magnetite and Alloy 600 and determined that they have the same sign of charge in high temperature alkaline water resulting in a repulsive force between the particle and surface, thereby significantly limiting deposition (Turner & Klimas, 2000). Therefore, it is possible that the magnetite particles formed in these experiments in subcritical water experienced repulsion due to the electric double layer repulsion.  On the other hand, surface attachment behaviour can be significantly different in supercritical water. The electric double layer force was calculated by Ghosh et al. (2006) as a function of temperature in subcritical and supercritical water at 25MPa. These forces were found to decrease significantly in supercritical water due to its low dielectric constant. On the other hand, Van der Waals attraction potential was shown to increase as temperature increased in  107  subcritical water, and further increase in supercritical water (Ghosh et al., 2006). This would suggest that in supercritical water, double layer and Van der Waals forces should not limit deposit growth. Instead, the lower deposition would most likely be the result of the removal process. This may be particularly important for particulate fouling which has a weaker bond to the surface than crystallization fouling.  Another limitation to the model is that the nucleation and growth of the particles was assumed to be uniform and instantaneous. Agglomeration of particles was also not considered as it is difficult to model the interactions of particles in a complex system (Bott, 1995). Although the formation of nanometer sized particles can be very fast (Adschiri et al., 1992), SEM images along the test section has shown that micron-sized particles require time to grow. Therefore, better models could be produced by including crystal growth along the test section.  108  7 Conclusion In this study, fouling and transport of magnetite particles in supercritical water were examined experimentally and through computer simulation models. Experimental work was conducted with the objective of utilizing several online and offline techniques for characterizing the deposition and transport phenomena under different heat flux, temperature ranges, and pH conditions. To form simulated corrosion products, a hydrothermal synthesis technique was adapted for producing magnetite particles in deaerated supercritical water using ferrous chloride and ferrous sulfate as precursors. EDX analysis of the particles collected on the filter clearly show an iron oxide composition and XRD of the tube and filter deposits confirm the formation of magnetite.  An online monitoring method utilizing thermal resistance properties of magnetite was implemented in the test section to infer the deposition thickness on the tube wall. Deposits led to an increase in outer wall temperature under supercritical conditions during the injection of ferrous chloride with no pH adjustment and ferrous sulfate at high pH. It is speculated that the mixed deposition of magnetite and precipitated ferrous precursor was the source of very high fouling rates in supercritical water. For high pH ferrous sulfate condition, deposition-removal cycles were observed at subcritical temperatures which are believed to be only magnetite particulate fouling. These cycles continued throughout the injection with complete removal occurring when deposit thickness approached 42 µm. It was observed that removal during cool down was particularly important for high pH experiments where sub-micron sized particles of 100 - 200 nm diameter were observed.  109  Using ultrasound and acid wash cleaning procedures, a qualitative method for determining deposit strength was developed. Results suggest that under supercritical conditions, the particulate deposit had a stronger adhesion to the tube and required an aggressive cleaning procedure to completely remove the deposit. It is possible that the precipitation of ferrous species between the depositing particle and surface was responsible for the stronger bond. In addition, SEM imaging of the test section showed the morphology of the deposit varied along the tube length and corresponded to the particles found on the filters.  All four techniques individually provide unique data on the particles while the combination of the techniques give an overall understanding of deposition/transport in the system. The online thermal resistance monitoring provided valuable information about deposition and removal cycles, while particle size and adhesive strength could only be determined from offline methods. In the future, high temperature high pressure filters could be implemented in full scale SCW systems which could provide valuable information on the effects of water chemistry on corrosion product transport. On the other hand, destructive testing techniques such as surface SEM and cleaning methods would likely be useful for small test setup experiments such as the one demonstrated in this study.  Simulation using heat and mass transport equations produced comparable predictions to experiments for the deposition thickness in the test section. In all cases, the simulation overestimated the deposit thickness, particularly for the high pH experiment where high removal was observed experimentally. The results suggest that both surface attachment and removal rates should be included in the simulation model which would effectively reduce the deposition  110  thickness. In supercritical conditions, the effects of surface attachment forces become less significant and removal is likely to be the dominant limitation to deposition. However, difficulty in modeling both removal and surface attachment stems from the lack of fundamental theory for predicting their coefficients, can vary significantly depending on environmental conditions.  Finally, the simulation was adapted to predict fouling rates in a hypothetical CANDU SCWR. The design parameters of a current ACR-700 fuel bundle along with proposed SCWR parameters provided a baseline for the simulation. Thermodynamic solubility predictions using HKF models from literature were used to predict the concentration of magnetite particles which would form in the reactor core. These simulations (assuming mass-transfer limited deposition) suggested that fouling may lead to an increase of fuel cladding-surface temperatures of up to 23.9⁰C after one year of operation assuming the formation of submicron particles. At 4 m down from the inlet, the temperature rise is expected to be 19.6ºC, raising the peak temperature of one section from 802.2ºC to 823.7ºC. This temperature increase in the fuel cladding can potentially be a significant problem for long term operation of a CANDU SCWR, and therefore requires further research.  111  8 Recommendations This study showed that fouling and transport of corrosion products can be a significant issue for the development of a Generation IV CANDU SCWR. Therefore there is a need for developing new techniques and instruments for monitoring and characterizing deposition and transport, particularly for magnetite particles. It remains a challenging task to produce these instruments which can withstand the high temperature and pressure of supercritical water. The following provide recommendations for further development for the techniques presented in this study.   Use nickel alloys for test section tube material which will provide better resistance to hydrochloric acid and are immune to sulfuric acid, both a byproduct of the hydrolysis of ferrous chloride and sulfate respectively, making it more suitable for hydrothermal synthesis of magnetite and cleaning method analysis (Kritzer, 2004).    Experiments of longer duration would be useful for validating the online thermal resistance measurements in more realistic deposition rates. Thermocouple drift may become an issue when the thermocouples are exposed to high temperature for extended period of time.    Experiments of very long durations (>1000 hrs) could simulate deposition and transport of corrosion products without the need for artificially introduced ferrous species. Instead, the water could be saturated with ferrous species at 350°C which is then pumped into a test section for deposition. This would require a fully automated system and the system piping would need to be constructed from an iron free alloy or lined with ceramic or glass.  112  References Adschiri, T., Hakuta, Y., & Arai, K. (2000). Hydrothermal synthesis of metal oxide fine particles at supercritical conditions. Industrial & Engineering Chemistry Research, 39, 4901-4907. Adschiri, T., Kanazawa, K., & Arai, K. (1992). Rapid and continuous hydrothermal crystallization of metal oxide particles in supercritical water. Journal of American Ceramic Society, 4, 1019-1022. American Society for Metals. (1985). Metals Handbook Desk Edition. Ohio: American Society for Metals. Anthony, J. W., Bideaux, R. A., Bladh, K. W., & Nichols, M. C. (1997). Handbook of Mineralogy (Vol. 3). Chantilly, VA: Mineralogical Society of America. Arden, T. V. (1950). The solubility products of ferrous and ferrosic hydroxids. Journal of Chemical Society, 24, 882-885. Basset, M., McInerney, J., Arbeau, N., & Derek, L. H. (2000). The fouling of alloy-800 heat exchange surfaces by magnetite particles. The Canadian Journal of Chemical Engineering, 78, 40-52. Bastidas, J. M., Torres, C. L., Cano, E., & Polo, J. L. (2002). Influence of molybdenum on passivation of polarised stainless steels in a chloride environment. Corrosion Science, 44, 625-633. Bazargan, M., & Fraser, D. (2009). Heat transfer to supercritical water in a horizontal pipe: modeling, new empirical correlation, and comparison against experimental data. Journal of Heat Transfer, 131, 1-9. Bazargan, M., Fraser, D., & Chatoorgan, V. (2005). Effect of buoyancy of heat transfer in supercritical water flow in a horizontal round tube. Journal of Heat Transfer, 127, 897902.  113  Beal, S. K. (1970). Deposition of particles in turbulent flow on channel or pipe wall. Nuclear Science and Engineering, 40, 1-11. Bott, T. R. (1995). Fouling of Heat Exchangers. Netherlands: Elsevier Science. Burrill, K. A. (1977). Corrosion product transport in water-cooled nuclear reactors. Part I: Pressurized water operatoin. The Canadian Journal of Chemical Engineering, 55, 54-61. Burrill, K. A. (2000). Water chemistries and corrosion product transport in supercritical water in reactor heat transport systems. Water Chemistry of Nuclear Reactor Systems, 1, 357-363. Cengel, Y. A., & Cimbala, J. M. (2006). Fluid Mechanics. New York: McGraw-Hill. Chen, Q., & Ahmadi, G. (1997). Deposition of particles in turbulent pipe flow. Journal of Aerosol Science, 28(5), 789-796. Cleaver, J. W., & Yates, B. (1973). Mechanism of detachment of colloidal particles from a flat substrate in a turbulent flow. Journal of Colloid and Interface Science, 44, 464-474. Cook, W. G., & Fatoux, W. (2009). A CANDU-SCWR with a steam generator: Thermodynamic assessment and estimation of fouling rates. Paper presented at the 4th International Symposium on Supercritical Water-Cooled Reactors, Heidelberg. Cooper, K., Gupta, A., & Beaudoin, S. (2001). Simulation of the adhesion of particles to surfaces. Journal of Colloid and Interface Science, 234, 284-292. Crockett, H. M., & Horowitz, J. S. (2010). Erosion in nuclear piping systems. Journal of Pressure Vessel Technology, 132, 1-3. Dai, L., Yu, M., & Dai, Y. (2007). Nozzle passage aerodynamic design to reduce solid particle erosion of a supercritical steam turbine control stage. Wear, 262, 104-111. Electric Power Research Institute Inc. (2003). Deposition in boilers - Review of Soviet and Russian literature. Concord, CA. Epstein, N. (1987). General Thermal Fouling Models. Fouling Science and Technology.  114  Epstein, N. (1997). Elements of particle deposition onto nonporous solid surfaces parallel to suspension flows. Experimental Thermal and Fluid Science, 14, 323-334. Gao, X., Wu, X., Zhang, Z., Guan, H., & Han, E. (2007). Characterization of oxide films grown on 316L stainless steel exposed to H2O2-containing supercritical water. Journal of Supercritical Fluids, 42, 157-163. Generation IV International forum SCWR Committee. (2002). A Technology Roadmap for Generation IV Nuclear Energy Systems. Ghosh, S. K., Alargova, R. G., Deguchi, S., & Tsujii, K. (2006). Dispersion stability of colloids in sub- and supercritical water. Journal of Physical Chemistry, 110, 25901-25907. Guzonas, D., Tremaine, P., & Brosseau, F. (2009). Predicting activity transport in a supercritical water cooled pressure tube reactor. Paper presented at the 4th International Symposium on Supercritical Water-Cooled Reactors, Heidelberg. Helling, R. K., & Tester, J. W. (1988). Oxidation of simple compounds and mixtures in supercritical water: Carbon monoxide, ammonia, ethanol. Environmental Science & Technology, 22(11), 1319-1324. Holman, J. P. (2002). Heat Transfer (9 ed.). New York: McGraw-Hill. Hong, Y. (2007). Composite Fouling on Heat Exchanger Surfaces. New York: Nova Science Publishers Inc. International Atomic Energy Agency. (2010). Energy, Electricity and Nuclear Power Estimates for the Period up to 2050. Vienna, Austria: IAEA. Jones, D. A. (1996). Principles and Prevention of Corrosion. Upper Saddle River, NJ: PrenticeHall. Jones, R. L., Gilman, J. D., & Nelson, J. L. (1993). Controlling stress corrosion cracking in boiling water reactors. Nuclear Engineering and Design, 143, 111-123.  115  Khan, M. S. (2005). Deposition of sodium carbonate and sodium sulfate in supercritical water oxidation systems and its mitigation. PhD Thesis, University of British Columbia. Khartabil, H. (2009). SCWR: Overview. Paper presented at the Gen-IV International Forum, Paris, France. Kritzer, P. (2004). Corrosion in high-temperature and supercritical water and aqueous solutions: a review. The Journal of Supercritical Fluids, 29, 1-29. Li, C., Shan, J., & Leung, L. K. H. (2009). Subchannel analysis of CANDU-SCWR fuel. Progress in Nuclear Energy, 799-804. Lister, D. H. (1980). Corrosion products in power generating systems. New York: Hemisphere Publishing Corporation. Marrone, P. A., Cantwell, S. D., & Dalton, D. W. (2005). SCWO system designs for waste treatment: Application to chemical weapons destruction. Industrial & Engineering Chemistry Research, 44, 9030-9039. Masuyama, F. (2001). History of power plants and progress in heat resistance steels. ISIJ International, 612-625. McCabe, L. P., Sargent, G. A., & Conrad, H. (1985). Effect of microstructure on the erosion of steel by solid particles. Wear, 105, 257-277. Miller, D. G. (1982). Estimation of tracer diffusion coefficients of ions in aqueous solution. Livermore, California: Lawrence Livermore Laboratory. Müller-Steinhagen, H., Reif, F., Epstein, N., & Watkinson, A. P. (1988). Influence of operating conditions on particulate fouling. The Canadian Journal of Chemical Engineering, 66, 42-50. Naidin, M., Mokry, S., Baig, F., Gospodinov, Y., Zirn, U., Pioro, I., & Naterer, G. (2009). Thermal-design options for pressure-channel SCWRs with cogeneration of hydrogen. Journal of Engineering for Gas Turbines and Power, 131, 1-8.  116  Newson, I. H., Bott, T. R., & Hussain, C. I. (1983). Studies of magnetite deposition from a flowing suspension. Chemical Engineering Communications, 20, 335-353. Papavergos, P. G., & Hedley, A. B. (1984). Particle deposition behaviour from turbulent flows. Chemical Engineering Research & Design, 62, 275-295. Propp, A. W., Carleson, T. E., Wai, C. M., Taylor, P. R., Daehling, K. W., Huang, S., & AbdelLatif, M. (1996). Corrosion in Supercritical Fluids. Idaho Falls: Idaho National Engineering Laboratory. Schwarz, T. (2001). Heat transfer and fouling behaviour of Siemens PWR steam generators long-term operating experience. Experimental Thermal and Fluid Science, 25, 319-327. Shaw, R. W., Brill, T. B., Clifford, A. A., Eckert, C. A., & Franck, E. U. (1991). Supercritical water: A medium for chemistry. Chemical Engineering News, 69(51), 26-39. Shreir, L. L., Jarman, R. A., & Burstein, G. T. (1994). Corrosion (3rd Edition) Volumes 1-2. Oxford: Butterworth-Heinemann. Sinha, A. K. (1989). Ferrous Physical Metallurgy. Stoneham: Butterworth Publishers. Srisukvatananan, P., Lister, D. H., Svoboda, R., & Daucik, K. (2007). Assessment of the state of the art of sampling of corrosion products from water/steam cycles. PowerPlant Chemistry, 9(10), 613-626. Sue, K., Adschiri, T., & Arai, K. (2002). Predictive model for equilibrium constants of aqueous inorganic species at subcritical and supercritical conditions. Industrial & Engineering Chemistry Research, 41, 3298-3306. Sugimoto, T., & Matijevic, E. (1980). Formation of uniform spherical magnetite particles by crystallization from ferrous hydroxide gels. 74(1), 227-243. Sweeton, F. H., & Baes Jr., C. F. (1970). Th solubility of magnetite and hydrolysis of ferrous ion in aqueous solutions at elevated temperatures. Journal of Chemical Thermodynamics, 2, 479-500.  117  Swenson, H. S., Carver, J. R., & Kakarala, C. R. (1965). Heat transfer to supercritical water in smooth-bore tubes. Journal of Heat Transfer, 477-484. Teshima, P. (1997). Fouling rates from a sodium sulphate - water solution in supercritical water oxidation reactors. MASc Thesis, University of British Columbia. Thomas, D., & Grigull, U. (1974). Experimental investigation of the deposition of suspended magnetite from the fluid flow in steam generating boiler tubes. Brennst-Warme-Kraft, 26, 109-117. Torgerson, D. F., Shalaby, B. A., & Pang, S. (2006). CANDU technology for Generation III+ and IV reactors. Nuclear Engineering and Design, 236, 1565-1572. Tremaine, P. R., & LeBlanc, J. C. (1980). The solubility of magnetite and the hydrolysis and oxidation of Fe2+ in water to 300C. Journal of Solution Chemistry, 9(6), 415-442. Turner, C. W. (1993). Rates of particle deposition from aqueous suspensions in turbulent flow: A comparison of theory with experiment. Chemical Engineering Science, 48, 2189-2195. Turner, C. W., & Klimas, S. J. (2000). Deposition of magnetite particles from flowing suspensions under flow-boiling and single-phase forced-convective heat transfer. The Canadian Journal of Chemical Engineering, 78, 1065-1075. Turner, C. W., Lister, D. H., & Smith, D. W. (1990). The deposition and removal of sub-micron particles of magnetite at the surface of alloy 800. Paper presented at the Steam Generator and Heat Exchanger Conference, Toronto, Canada. Viswanathan, R., Henry, J. F., Tanzosh, J., Stanko, G., Shingledecker, J., Vitalis, B., & Purgert, R. (2005). U.S. program on materials technology for ultra-supercritical coal power plants. Journal of Materials Engineering and Performance, 14(3), 281-292. Was, G. S., Ampornrat, P., Gupta, G., Teysseyre, S., West, E. A., Allen, T. R., . . . Pister, C. (2007). Corrosion and stress corrosion cracking in supercritical water. Journal of Nuclear Materials, 176-201.  118  Was, G. S., Teysseyre, S., & Jiao, Z. (2006). Corrosion of austenitic alloys in supercritical water. Corrosion, 62(11). Weingartner, H., & Franck, E. U. (2005). Supercritical water as a solvent. Angewandte Chemie International Edition, 44(18), 2672-2692. Willard, H. H., Merritt, L. L. J., & Dean, J. A. (1969). Instrumental Methods of Analysis (4 ed.). London: D. Van Nostrand Company Inc. Yamagata, K., Nishikawa, K., Hasegawa, S., Fujii, T., & Yoshida, S. (1971). Forced convective heat transfer to supercritical water flowing in tubes. International Journal of Heat and Mass Transfer, 15, 2575-2593. Yeon, J., Jung, Y., & Pyun, S. (2006). Deposition behaviour of corrosion products on the zircaloy heat transfer surface. Journal of Nuclear Materials, 354, 163 - 170. Zhang, L., Zhu, F., & Tang, R. (2009). Corrosion mechanisms of candidate structural materials for supercritical water-cooled reactor. Front. Energy Power Engineering China, 3(2), 233-240. Zhao, L., Zhang, H., Tang, J., Song, S., & Cao, F. (2009). Fabrication and characterization of uniform Fe3O4 octahedral microcrystals. Materials Letters, 63, 307-309.  119  Appendices Appendix A: MATLAB Code %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % main.m % % This is the main program to run to calculate heat transfer and deposition % Written by Keigo Karakama % Date created: September 25, 2010 % Last modified: October , 2011 % Clear everything from previous programs clc clear all %------------------------------------------------------------------------% Section 1: User defined inputs flowrate Ti To P di do tlength runtime_min Cb(1)  = = = = = = = = =  3.50; 350; 400; 237; 0.069*0.0254; 0.125*0.0254; 1.75; 40; 500*10^-6;  dp corr flowrate_pump2  = 0.1 *10^-6; = 2; = 0.5;  % % % % % % % % % %  Total flow rate (Pump 1 and 2) Temp. of bulk fluid going in [deg C] Temp. of bulk fluid going out [deg C] Pressure [bar] Inside tube dia. of test section [m] Outside tube dia. of test section [m] Length of tube [m] Run time of injection [min] Initial concentration of suspended... magnetite in solution [kg/L] % Diameter of particle [m] % Correlation to use for heat transfer % Injection pump flow rate [L/hr]  %--------------------------------------------------------------------------% Section 2: Runtime Variables m Nlength dt dx x h_upstream h_downstream delta_h Q q hflux  = = = = = = = = = = =  flowrate/3600; % Flow rate in [kg/s] 100; % Number of length steps runtime_min*60 % Time step [s] tlength/Nlength; % Increment [m] 0:dx:tlength; % Array XSteam('h_pT',P,Ti); XSteam('h_pT',P,To); (h_downstream - h_upstream)/Nlength; (h_downstream - h_upstream)*m; % [kW] Q*1000/(pi * di); % [W/m] Q*1000/(pi * di * tlength); % [W/m^2]  %-------------------------------------------------------------------------% Section 3: Constants rho_oxide K316L  = 5200; = 20.2;  % Density of magnetite [kg/m^3] % Thermal conductivity of 316L [W/mK]  120  % Other Variables % KFe3O4 Thermal conductivity of Fe3O4 [W/mK] % Tw1 Temperature at the bulk/surface interface % Tw2 Temperature at the point between the oxide and alloy % Tw3 Outside wall temperature and is measured experimentally %-------------------------------------------------------------------------% Section 4: Set Initial Values count = 0; % Set initial count of loop to zero Tb(1) = Ti; % Temperature of bulk fluid dmdt(1) = 0; % Deposition mass flux dm(1) = 0; % Deposition mass thickness = zeros(1,Nlength+1); [buo(1),Re(1),rhoB(1),Tw1(1),hconv(1)] ... = fluidproperties(m,Tb(1),di,P,hflux,corr); Tw2(1) = Tw1(1); % Temperature inside wall Tw3(1) = twecalc(Tw2(1),hflux,K316L,di,do); h(1) = h_upstream; %-------------------------------------------------------------------------% Section 5: Main part of the code for j = 2:Nlength+1, % Track and print to display count = 1 + count % Calculate fluid properties of next step h(j)=h(j-1) + delta_h; % Enthalpy at position x Tb(j) = XSteam('T_ph',P,h(j)); % Bulk temperature at position x [buo(j),Re(j),rhoB(j),Tw1(j),hconv(j)]... = fluidproperties(m,Tb(j),di,P,hflux,corr); % For preformed particle transport, saturation is negligible Csat(j) = 0; % Calculate bulk concentration calculated from previous deposition Cb(j) = Cb(j-1) - dmdt(j-1)/m; % If bulk concentration is not zero, then particles are in the fluid if (Cb(j) > Csat(j)) [Kt(j),TP(j), LN(j),viscouslayer(j)] = depositionvelocity(m,... rho_oxide,P,Tb(j),dp,di,thickness(j)); dmdt(j) = depositionfun(Kt(j),dx,di,rho_oxide,Cb(j),... Csat(j)); dm(j) = dmdt(j)*dt; rfoul(j) = thicknessfun(dmdt(j),rho_oxide,di/2,dx,dt); % If bulk concentration is zero, set values to default value of 0 else Kt(j) = 0; TP(j) = 0; LN(j) = 0; viscouslayer(j) = 0; dmdt(j) = 0;  121  dm(j) rfoul(j)  = 0; = 0;  end % Calculate the cumulative thickness of the deposit thickness(j) = rfoul(j) + thickness(j); % Calculate the thermal conductivity of magnetite at the temperature KFe3O4(j) = 4.133 - 0.852*10^-2*Tw1(j) + 0.757*10^-5*Tw1(j); % Calculate the temperature at the surface interfaces Tw2(j) = twocalc(Tw1(j),hflux,KFe3O4(j),... (di-2*thickness(j)),di); Tw3(j) = twecalc(Tw2(j),hflux,K316L,di,do); end %-------------------------------------------------------------------------% Section 6: Plotting xnew = dx:dx:tlength; %For graphs which does not need value at x=0 for k=1:Nlength, yfoul(k) = thickness(k+1)*1000000; end % Plot of result summary figure(1) subplot(2,2,1) plot(xnew,yfoul) xlabel('Length [m]') ylabel('Thickness [\mum]') title('Thickness vs Length') subplot(2,2,2) semilogy(x,TP,x,0.2,x,20) xlabel('Length [m]') ylabel('tp*') title('tp* vs Length') subplot(2,2,3) plot(x,Tb) xlabel('Length [m]') ylabel('Temperature [(Khan)C]') title('Bulk Temperature vs Length') subplot(2,2,4) plot(x,Cb*1000000) xlabel('Length [m]') ylabel('Concentration [mg/kg]') title('Concentration vs Length') % Plot of temperature profile figure(2) plot(x,Tb,'k.-',x,Tw1,'k:',x,Tw2,'k--',x,Tw3,'k') legend('Bulk fluid temperature','Oxide temperature', ... 'Tube inside wall temperature','Tube outside wall temperature'); xlabel('Length (m)')  122  ylabel('Temperature (C)') % Plot of fouling thickness vs. distance figure(3) plot(xnew,yfoul) xlabel('Distance [m]') ylabel('Magnetite thickness [\mum]') % Plot of Reynolds vs. distance figure(4) plot(x,Re) xlabel('Distance [m]') ylabel('Re') % Plot of heat transfer coefficient vs. distance figure(5) plot(x,hconv) xlabel('Distance [m]') ylabel('Heat transfer coefficient') % Plot of Gr/Grth vs. distance for buoyancy figure(6) plot(x,buo,'k') xlabel('Distance [m]') ylabel('Gr/Grth') %-------------------------------------------------------------------------% Section 6: Reporting to MATLAB display thickp = thickness'*1000000; xp = x'; Bulkfluid = Cb(Nlength+1); section_mass = zeros(10,1); count1 = 1; count2 = 1; for p=1:Nlength, section_mass(count1) = section_mass(count1)+dm(p); if count2 == (Nlength/10) count1 = count1+1; count2 = 1; else count2 = count2+1; end end section_mass_mg total_mass_mg deposited_mass_mg mass_to_filter_mg deposited_percentage concentration_tank2_mgL total_mass_g transpose(thickness); hflux  = = = = = = =  section_mass*1000000 Cb(1)*flowrate*runtime_min*1000000/60 sum(section_mass_mg) total_mass_mg - deposited_mass_mg deposited_mass_mg*100/total_mass_mg Cb(1)*(flowrate/flowrate_pump2)*1000000 Cb(1)*flowrate*runtime_min*1000/60  123  Functions %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Buoyancy Effects Calculation % Bouyancy Calculations based from Petukhov, evaluated by Bazargan % Written by Keigo Karakama % Date created: April 18, 2011 % Last modified: July 19, 2011 function [A,Grth,Grq] = bouyancy(Tw,Tb,Reb,d,P,m,hflux) % % % % % % % %  Grth Grq Pr Hw,Hb Tw,Tb mu k hflux  g Tf rhoF rhoB rhoW Hw Hb k mu  = = = = = = = =  Grashof number threshold Grashof number heat-flux-related Prandtl number Enthalpy at wall and bulk respectively Temperature at wall and bulk respectively bulk dynamic viscosity bulk thermal conductivity heat flux  = = = = = = = = =  9.81; (Tw + Tb)/2; XSteam('rho_pT' XSteam('rho_pT' XSteam('rho_pT' XSteam('h_pT' XSteam('h_pT' XSteam('tc_pT' XSteam('my_pT'  ,P,Tf); ,P,Tb); ,P,Tw); ,P,Tw); ,P,Tb); ,P,Tb); ,P,Tb);  Pr = (Hw-Hb)*1000*(mu/k)/(Tw-Tb); Refun_pipe(m,d,mu); B = (1/rhoF)*(rhoB - rhoW)/(Tw-Tb); Grth = (3*10^-5)*(Reb^2.75)*(Pr^0.5)*(1+2.4*(Reb^-0.125)*(Pr^(2/3)-1)); Grq = g*B*hflux*d^4/((mu/rhoB)^2*k); A = Grq/Grth; end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Deposition Velocity % Written by Keigo Karakama % Date created: January 9, 2011 % Last modified: January 20, 2011 function dmdt = depositionfun(Kt,dx,di,rho_oxide,Cb,Csat); % % % % % %  Kt = di = dx = rhoB Cb = Csat  mass transport deposition velocity [m/s] inner diameter of tube [m] step size [m] = bulk fluid density [kg/m^3] bulk concentration [kg Fe3O4/kg H2O] = Saturation concentration [kg Fe3O4/kg H2O]  124  dmdt = rho_oxide*Kt*(Cb-Csat)*pi*di*dx; end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Deposition Velocity % Written by Keigo Karakama % Date created: January 9, 2011 % Last modified: January 10, 2011 function [Kd TP LN delta] = depositionvelocity(m,rhop,P,Tb,dp,di,thickness) % % % % % %  m rhop P Tb dp di  rhof= mu = Re = eps = f = u = U = D = TP = Sc =  = = = = = =  mass flow rate [kg/s] density of particle [kg/m3] pressure [bar] temperature of bulk fluid [C] diameter of particle [m] diameter of inner wall of tube [m]  XSteam('rho_pT',P,Tb); XSteam('my_pT',P,Tb); Refun_pipe(m,di,mu); 0.000002 + thickness; frictionfun(eps,di,Re); 4*m/(rhof*pi*di^2); u*sqrt(f/8); diffusioncoefficient(Tb,mu,dp); rhop*(rhof*U*dp/mu)^2/(18*rhof); Scfun(mu,rhof,D);  % Wood, Fan and Ahmadi Equation if TP < 0.2 KD = (4.5*10^-4)*TP^2+0.057*Sc^(-2/3); LN = 1; elseif (((TP > 0.2)||(TP == 0.2))&&((TP < 20)||(TP == 20))) KD = (3.5*10^-4)*TP^2; LN = 2; elseif TP > 20 KD = 0.18; LN = 3; end % Calculate the deposition velocity [m/s] Kd = KD*U; % Calculate the viscous sublayer distance [m] delta = 5*mu/(rhof*U); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Diffusion Coefficient % Written by Keigo Karakama % Date created: January 9, 2011  125  % Last modified: January 9, 2011 function D = diffussioncoefficient(T,mu,dp) % % % %  kB T mu dp  = = = =  Boltzmann constant [J/K] temperature [K] dynamic viscosity [Pa.s] diameter of particle [m]  kB = 1.38065*10^-23; Tk = T + 273; D = kB*Tk/(3*pi*mu*dp); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Fluid/thermal Properties Function % Written by Keigo Karakama % Date created: August 23, 2010 % Last modified: October 9, 2011 function [A,Reb,rhoB,Tw,hconv] = ... fluidproperties(m,Tb,di,P,hflux,select) % % % % % % %  m Tw Tb di P hflux k  = = = = = = =  mass flow rate [kg/s] wall temperature [C] bulk fluid temperature [C] inside diameter [m] pressure [bar] heat flux [W/m^2] thermal conductivity of fluid [W/m.K]  % Calculate bulk fluid properties rhoB = XSteam('rho_pT' ,P,Tb); % First estimate to Tw Tw = Tb + 1; Tw_old = Tb + 2; % Determine pseudocritical temperature Tpc = pseudocritical(P); % Correlation by Swenson et al. if select == 1 while (abs(Tw-Tw_old) > 0.2) % Calculate fluid properties Hw = XSteam('h_pT' ,P,Tw)*1000;% Hb = XSteam('h_pT' ,P,Tb)*1000;% k = XSteam('tc_pT' ,P,Tw); % rhoW = XSteam('rho_pT' ,P,Tw); % viscosity = XSteam('my_pT' ,P,Tw); % cp = (Hw - Hb)/(Tw-Tb); %  [J/kg] [J/kg] [W/m.K] [kg/m^3] [Pa.s] [J/kgK]  % Calculate non-dimensional numbers Re = Refun_pipe(m,di,viscosity);  126  Pr Nu hconv  = Prfun(viscosity,cp,k); = Nufun(Re,Pr,rhoB,rhoW,Tb,P); = Nu*k/di;  % Store the old wall temperature and calculate a new one Tw_old = Tw; Tw = Twfun(hflux,Nu,k,Tb,di); Tw = (Tw+Tw_old)/2; end % Correlation by Bazargan and Fraser elseif select == 2; % Calculate enthalpy of pseudocritical and bulk temperatures Hpc = XSteam('h_pT' ,P,Tpc); Hb = XSteam('h_pT' ,P,Tb); % Initial guess of Tw Tw = Tb + 1; Tw_old = Tb + 2; while (abs(Tw-Tw_old) > 0.2) Hw = XSteam('h_pT' ,P,Tw); Izone = (Hpc - Hb) / (Hpc - Hw); if ((Izone > -0.9)&&(Izone < 1)) Hfac = (1-Izone)*(Hpc)/(1.9*Hw); LM = 1; else Hfac = 0; LM = 2; end Href Tref cp k rhoW viscosity Re Pr  = = = = = = = =  Hb + Hfac*(Hw - Hb); XSteam('T_ph' ,P,Href); XSteam('Cp_pT' ,P,Tref) * 1000; %J/kg.K XSteam('tc_pT' ,P,Tref); % W/mK XSteam('rho_pT' ,P,Tref); % kg/m3 XSteam('my_pT' ,P,Tref); % Pa.s Refun_pipe(m,di,viscosity); Prfun(viscosity,cp,k);  % Dittus Boelter Correlation Nu = 0.023*Re^0.8*Pr^0.4; hconv Tw_old Tw Tw end  = = = =  Nu*k/di; Tw; Twfun(hflux,Nu,k,Tb,di); (Tw+Tw_old)/2;  end end % Calculate bouyancy effects for this temperature and pressure viscB = XSteam('my_pT' ,P,Tb); % Pa.s Reb = Refun_pipe(m,di,viscB);  127  [A,Grth,Grq]  = bouyancy(Tw,Tb,Reb,di,P,m,hflux);  end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Friction Factor Calculation % Written by Keigo Karakama % Date created: January 12, 2011 % Last modified: January 12, 2011 % % Note: Uses explicit relation developed by S.E.Haaland function f = frictionfun(eps,di,Re) % Re = Reynolds number % di = diameter of inside of tube [m] % eps = roughness value of tube [m] f = (-1.8*log10((6.9/Re)+(eps/(3.7*di))^1.11))^-2; end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Nusselt number function % Written by Keigo Karakama % Created: November 28, 2010 % Modified: December 9, 2010 function Nu = Nufun(Re,Pr,rhoW,rhoB,Tb,P) % % % % %  Re = Reynolds number Pr = Prandtl number rhoW, rhoB = fluid density at wall and bulk respectively [kg/m^3] Tb = Bulk fluid temperature [C] P = Pressure [bar]  Tcritical = 374; Pcritical = 221; if ((Tb > Tcritical)&&(P > Pcritical)) % Swenson et al. equation Nu = 0.00459*Re^0.923*Pr^0.613*(rhoW/rhoB)^.231; else % Dittus and Boelter equation Nu = 0.023*Re^0.8*Pr^0.4; end end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Prandtl Function % Written by Keigo Karakama % Date created: August 23, 2010 % Last modified: August 23, 2010  128  function prandtl = Prfun(viscosity,cp,k) % viscosity % cp % k  = viscosity [Pa.s] = specific isobaric heat capacity [J/kg.K] = thermal conductivity [W/mK]  prandtl = cp*viscosity/k; end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Reynolds Pipe Function % Written by Keigo Karakama % Date created: August 23, 2010 % Last modified: December 9, 2010 function reynolds = Refun_pipe(m,D,viscosity) % % % % % %  D = inner tube diameter [m] m = mass [kg/s] viscosity = viscosity [Pa.s] Laminar Re ~< 2300 Transitional 2300 < Re < 4000 Turbulent Re > 4000  reynolds = 4*m/(pi*D*viscosity); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Schmidt Number = viscous diffusion rate / mass diffusion rate % Written by Keigo Karakama % Date created: January 9, 2011 % Last modified: January 9, 2011 function Sc = Scfun(mu,rho,D) % mu = dynamic viscosity [Pa.s] % rho = density [kg/m^3] % D = diffusion coefficient [m^2/s] Sc = mu/(rho*D); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Thickness of deposit calculation % Written by Keigo Karakama % Date created: December 1, 2010 % Last modified: January 24, 2011 % Equation from Holman (see reference) function t = thicknessfun(m_oxide,rho_oxide,r2,dx,dt)  129  % m_oxide = mass of deposit [kg] % rho_oxide = density of oxide [kg] % r2 = radius [m] t = r2 - sqrt(r2^2 - m_oxide*dt/(rho_oxide*pi*dx)); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % External Wall Temperature % Program to calculate the outside wall temperature with internal heat % Written by Keigo Karakama % Date created: December 1, 2010 % Last modified: December 1, 2010 function To = twecalc(Ti,hflux,k,di,do) % % % % % %  Ti To hflux k di do  = = = = = =  Inside wall temperature [deg C] Outside wall temperature [deg C] Heat flux [W/m2] Thermal conductivity of solid [W/mK] Inside diameter [m] Outside diamter [m]  A = di/do; To = Ti + hflux*(di/2)/(2*k)*((A^2-log(A^2)-1)/(1-A^2)); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%% % Oxide/Wall temperature % Written by Keigo Karakama % Date created: December 1, 2010 % Last modified: January 24, 2011 % Equation from Holman (see reference) function To = twocalc(Ti,hflux,k,di,do) % % % % % %  Ti = Inside wall temperature [C] To = Outside wall temperature [C] hflux = Heat flux [W/m2] k = Thermal Conductivity of solid [W/mK] di = Inside diameter [m] do = Outside diamter [m]  To = Ti + hflux*do*log(do/di)/(2*k); end %%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%%  130  Appendix B: Filter Parts Drawings  131  132  Appendix C: Results of Experiments C. 1. Reference Condition, Ferrous Chloride Table C-1: Summary table of experiment #1  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 19-Apr-11  Comments  FeCl2∙4H2O 23.7 5.48 3.00 0.54 3.54 0.03 3 42 1.8 15 350 400 102 11 Yes Yes Yes Yes Yes Yes  (1)  Comments: (1) Flow rates of the secondary line were found to be significantly lower than originally set and therefore there was very little filter deposits.  133  Injection  Figure C-1: Temperature measurements of test section vs. time for reference condition with ferrous chloride precursor. (Experiment #1)  Figure C-2: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for reference condition with ferrous chloride precursor. Low flow rate in secondary line resulting in very few particles collected. (Experiment #1) 134  Figure C-3: SEM photograph of particles collected on LTHP filter, 1.8k magnification, for reference condition with ferrous chloride precursor. (Experiment #1)  Figure C-4: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for reference condition with ferrous chloride precursor. Low flow rate in secondary line resulting in very few particles collected. (Experiment #1)  135  (a)  (b)  (c)  (d) Figure C-5: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #1)  136  Figure C-6: Deposit thickness on test section for reference condition with ferrous chloride precursor. (Experiment #1)  Figure C-7: Comparison between simulation and experimental for reference condition ferrous chloride. Simulation was run with an average particle size = 4 µm, magnetite concentration = 200 mg/L. 137  C. 2. Reference Condition, Ferrous Sulfate Table C-2: Summary table of experiment #2  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 24-Jun-11  Comments  FeSO4·7H2O 23.9 4.97 3.05 0.49 3.54 0.21 3.5 39 1.8 15 350 397 97 11 Yes Yes Yes Yes Yes Yes  138  Injection  Figure C-8: Temperature measurements of test section vs. time, temperature fluctuations were found in both the inlet and outlet bulk temperatures which may be attributed to the inconsistent flow. (Experiment #2)  Figure C-9: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for reference condition with ferrous sulfate. (Experiment #2) 139  Figure C-10: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for reference condition with ferrous sulfate. (Experiment #2)  Figure C-11: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for reference condition with ferrous sulfate. (Experiment #2)  140  Figure C-12: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #2)  141  Figure C-13: Deposit thickness on test section for reference condition with ferrous sulfate precursor. (Experiment #2)  Figure C-14: Comparison between simulation and experimental for reference condition ferrous sulfate. Simulation was run with an average particle size = 1 µm, magnetite concentration = 190 mg/L. 142  C. 3. Low Heat Flux, Ferrous Chloride Table C-3: Summary table of experiment #3  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 30-Mar-11 FeCl2∙4H2O 22.4 5.31 3.06 0.54 3.60 0.21 3 44 1 10 372 376 24 10 Yes Yes Yes No Yes Yes  Comments  (1)  (2)  Comments: (1) Experiment used a 1 m long test section and lower line voltage. (2) LTLP filter had not been added yet to the system for this experiment.  143  Injection  Figure C-15: Temperature measurements of test section vs. time for low heat flux ferrous chloride precursor. (Experiment #3)  Figure C-16: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for low heat flux with ferrous chloride precursor. (Experiment #3)  144  Figure C-17: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for low heat flux with ferrous chloride precursor. (Experiment #3)  Image Not Available (See comments in Summary)  Figure C-18: SEM photograph of particles collected on LTLP filter, - magnification, for low heat flux with ferrous chloride precursor. (Experiment #3)  145  (a)  (b)  (c) Figure C-19: SEM photograph of magnetite deposit on test section tube surface, (a) 0.10 m, (b) 0.30 m, and (c) 0.90 m location. (Experiment #3)  146  Data Not Available (See comments in Summary)  Figure C-20: Deposit thickness on test section for low heat flux ferrous chloride precursor. (Experiment #3)  147  C. 4. Low Heat Flux, Ferrous Sulfate Table C-4: Summary table of experiment #4  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 31-Aug-11  Comments  FeSO4·7H2O 23.0 5.04 3.03 0.51 3.54 0.20 3.5 40 1.8 15 372 376 13 11 Yes Yes Yes Yes No Yes  148  Injection  Figure C-21: Temperature measurements of test section vs. time for low heat flux with ferrous sulfate precursor. (Experiment #4)  Figure C-22: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for low heat flux ferrous sulfate precursor (Experiment #4)  149  Figure C-23: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for low heat flux ferrous sulfate precursor. (Experiment #4)  Figure C-24: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for low heat flux ferrous sulfate precursor. (Experiment #4)  150  (a)  (b)  (c)  (d) Figure C-25: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #4)  151  Figure C-26: Deposit thickness on test section for low heat flux ferrous sulfate precursor. (Experiment #4)  Figure C-27: Comparison between simulation and experimental for low heat flux, ferrous sulfate. Simulation was run with an average particle size = 1 µm, magnetite concentration = 108 mg/L. 152  C. 5. No Heat Flux, Ferrous Sulfate Table C-5: Summary table of experiment #5  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 20-Sept-11  Comments  FeSO4·7H2O 23.3 5.32 2.97 0.53 3.50 0.21 3.5 40 1.8 15 384 383 11 No Yes Yes Yes No Yes  (1)  Comments: (1) Because there was no net heat flux in this experiment, thermal resistance measurements on the test section were not taken.  153  Image Not Available (See comments in Summary)  Figure C-28: Temperature measurements of test section vs. time for no heat flux with ferrous sulfate precursor. (Experiment #5)  Figure C-29: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for no heat flux with ferrous sulfate precursor. (Experiment #5)  154  Figure C-30: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for no heat flux with ferrous sulfate precursor. (Experiment #5)  Figure C-31: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for no heat flux with ferrous sulfate precursor. (Experiment #5)  155  (a)  (b)  (c)  (d) Figure C-32: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #5)  156  Figure C-33: Deposit thickness on test section for no heat flux with ferrous sulfate precursor. (Experiment #5)  Figure C-34: Comparison between simulation and experimental for no heat flux with ferrous sulfate precursor. Simulation was run with an average particle size = 2 µm, magnetite concentration = 89 mg/L. 157  C. 6. High pH, Ferrous Sulfate Table C-6: Summary table of experiment #6  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 03-Aug-11  Comments  FeSO4·7H2O 23.8 5.11 2.99 0.51 3.50 0.23 9 NaOH 37 1.8 15 350 396 97 11 Yes Yes Yes Yes No Yes  158  Injection  Figure C-35: Temperature measurements of test section vs. time for high pH with ferrous sulfate precursor. (Experiment #6)  Figure C-36: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for high pH with ferrous sulfate precursor. (Experiment #6)  159  Figure C-37: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH with ferrous sulfate precursor. (Experiment #6)  Figure C-38: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for high pH with ferrous sulfate precursor. (Experiment #6)  160  (a)  (b)  (c)  (d) Figure C-39: SEM photograph of magnetite deposit on test section tube surface, (a) 0.15 m, (b) 0.60 m, (c) 1.20 m, and (d) 1.65 m location. (Experiment #6)  161  Figure C-40: Deposit thickness on test section for high pH with ferrous sulfate precursor. (Experiment #6)  Figure C-41: Comparison between simulation and experimental for high pH, ferrous sulfate. Simulation was run with an average particle size 0.15 µm, magnetite concentration = 321 mg/L. 162  C. 7. High pH & Subcritical, Ferrous Sulfate Table C-7: Summary table of experiment #7  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 22-Aug-11  Comments  FeSO4·7H2O 23.7 5.17 2.95 0.51 3.47 0.27 9 NaOH 40 1.8 15 200 370 92 11 Yes Yes No No No Yes  (1) (1)  Comments: (1) Test section was not disassembled after experiment.  163  Injection  Figure C-42: Temperature measurements of test section vs. time for high pH & subcritical with ferrous sulfate precursor. (Experiment #7)  Figure C-43: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for high pH & subcritical with ferrous sulfate precursor. (Experiment #7)  164  Figure C-44: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for high pH & subcritical with ferrous sulfate precursor. (Experiment #7)  Figure C-45: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for high pH & subcritical with ferrous sulfate precursor. (Experiment #7) 165  C. 8. Low Concentration, Ferrous Chloride Table C-8: Summary table of experiment #8  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 26-May-11  Comments  FeCl2∙4H2O 23.7 0.56 2.95 0.54 3.50 0.24 3 400 1.8 15 350 400 96 11 Yes Yes No No No No  (1) (1)  Comments (1) Test section was not disassembled after experiment.  166  Injection  Figure C-46: Temperature measurements of test section vs. time for low concentration with ferrous chloride precursor. (Experiment #8)  Figure C-47: SEM photograph of particles collected on HTHP filter, 2.0k magnification, for low concentration with ferrous chloride precursor. (Experiment #8)  167  Figure C-48: SEM photograph of particles collected on LTHP filter, 2.0k magnification, for low concentration with ferrous chloride precursor. (Experiment #8)  Figure C-49: SEM photograph of particles collected on LTLP filter, 2.0k magnification, for low concentration with ferrous chloride precursor. (Experiment #8)  168  C. 9. Blank Run Table C-9: Summary table of blank run  Main Parameters  Test Section  Analysis  Summary Date of Experiment Precursor Average Pressure Concentration Pump 1 Flow Rate Pump 2 Flow Rate Total Flow Rate Secondary Line Flow Rate pH pH Base Injection Time Length of Test Section Thermocouple Spacing Bulk Fluid Temperature Inlet Bulk Fluid Temperature Outlet Heat Flux In (based on enthalpy) # of Thermocouples (test section) Thermal Resistance Filters Deposit with SEM Deposit with Cleaning Methods XRD Effluent  Units  [MPa] [mmol/L] [L/hr] [L/hr] [L/hr] [L/hr]  [min] [m] [cm] [°C] [°C] [kW/m2]  Values 23-Jun-11  Comments  None 23.7 0.00 2.98 0.50 3.48 0.21 Neutral 40 1.8 15 351 399 99 11 Yes Yes No No No No  (1) (1)  Comments (1) Test section was not disassembled after experiment.  169  Figure C-50: Temperature measurements of test section vs. time for high pH & subcritical with ferrous sulfate precursor. (Experiment #9)  Figure C-51: SEM photograph of particles collected on HTHP filter, 2.0k magnification, blank run. (Experiment #9)  170  Figure C-52: SEM photograph of particles collected on LTHP filter, 5.0k magnification, blank run. (Experiment #9)  Figure C-53: SEM photograph of particles collected on LTLP filter, 2.0k magnification, blank run. (Experiment #9)  171  Appendix D: Filter Flow Rates Table D-1: Flow rate through filter  Flow Rate [L/hr]  2  Area of Filter Sample [cm ]  Filter  1  2  3  4  5  6  7  8  HTHP  0.034  0.21  0.21  0.20  0.21  0.23  0.27  0.24  LTHP  3.506  3.33  3.39  3.339  3.285  3.265  3.195  3.255  LTLP  0.034  0.21  -  0.201  0.21  0.23  0.27  0.24  HTHP  0.71  0.71  0.71  0.71  0.71  0.71  0.71  0.71  LTHP  2.85  2.85  2.85  2.85  2.85  2.85  2.85  2.85  LTLP  4.91  4.91  4.91  4.91  4.91  4.91  4.91  4.91  42  39  44  40  40  37  40  400  HTHP  0.033  0.191  0.216  0.188  0.196  0.199  0.252  2.244  LTHP  0.861  0.759  0.872  0.781  0.768  0.706  0.747  7.614  LTLP  0.005  0.028  -  0.027  0.029  0.029  0.037  0.326  LTHP/HTHP  25.8  4.0  4.0  4.2  3.9  3.6  3.0  3.4  Time of Injection [min] 2  Volume per Area [L/cm ]  Ratio  172  Appendix E: Atomic Absorption Spectroscopy The concentration of iron in each sample using AAS was determined from the calibration curve found in Figure E-1and Figure E-2 which produced equations E-1 and E-2 respectively.  Figure E-1: Calibration curve for iron concentration using AAS, nitric acid matrix  E-1  where C is the concentration [mg/L] and A is the absorbance.  173  Figure E-2: Calibration curve for iron concentration using AAS, nitric acid and copper sulfate matrix  E-2  The uncertainty in the thickness of the deposit was estimated from the uncertainty of several sources in the calculation including tube length, calibration curve, background concentration, and volume error.  Uncertainty from calibration curve: ±0.3 mg/L Uncertainty from background iron concentration = ±0.1 mg/L Uncertainty in volume = 5% Uncertainty in timing, heating, loss of iron in transfer = 10%  174  

Cite

Citation Scheme:

        

Citations by CSL (citeproc-js)

Usage Statistics

Share

Embed

Customize your widget with the following options, then copy and paste the code below into the HTML of your page to embed this item in your website.
                        
                            <div id="ubcOpenCollectionsWidgetDisplay">
                            <script id="ubcOpenCollectionsWidget"
                            src="{[{embed.src}]}"
                            data-item="{[{embed.item}]}"
                            data-collection="{[{embed.collection}]}"
                            data-metadata="{[{embed.showMetadata}]}"
                            data-width="{[{embed.width}]}"
                            async >
                            </script>
                            </div>
                        
                    
IIIF logo Our image viewer uses the IIIF 2.0 standard. To load this item in other compatible viewers, use this url:
http://iiif.library.ubc.ca/presentation/dsp.24.1-0072444/manifest

Comment

Related Items