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High Temperature Gas-cooled Reactors
 9780128210314

Table of contents :
Title-page_2021_High-Temperature-Gas-Cooled-Reactors
High Temperature Gas-cooled Reactors
Copyright_2021_High-Temperature-Gas-Cooled-Reactors
Copyright
Contents_2021_High-Temperature-Gas-Cooled-Reactors
Contents
List-of-contributors_2021_High-Temperature-Gas-Cooled-Reactors
List of contributors
About-the-authors_2021_High-Temperature-Gas-Cooled-Reactors
About the authors
Preface-of-JSME-Series-in-Thermal-and-Nuclear_2021_High-Temperature-Gas-Cool
Preface of JSME Series in Thermal and Nuclear Power Generation
Preface-to-Volume-5--High-Temperature-Gas-C_2021_High-Temperature-Gas-Cooled
Preface to Volume 5: High-Temperature Gas-Cooled Reactors
1---Overview-of-high-temperature-gas-coole_2021_High-Temperature-Gas-Cooled-
1 Overview of high temperature gas-cooled reactor
1.1 Features of high temperature gas-cooled reactor
1.1.1 Structure and materials
1.1.1.1 Fuel
1.1.1.2 Coolant
1.1.1.3 Moderator
1.1.2 Heat application
1.1.3 Safety
1.1.4 Adaptability to environment
1.2 History of research and development in world
1.3 History of research and development in Japan
References
2---Design-of-High-Temperature-Engineering-Te_2021_High-Temperature-Gas-Cool
2 Design of High Temperature Engineering Test Reactor (HTTR)
2.1 Overview of HTTR design features
2.1.1 Introduction
2.1.2 History and future plan of HTTR project
2.1.2.1 Evaluation of reactor performance
2.1.2.2 Safety demonstration test
2.1.2.3 Development of process heat application system
2.1.3 Major design features of HTTR
2.1.3.1 Reactor core
2.1.3.1.1 Core components
2.1.3.1.2 Reactor internals
2.1.3.1.3 Reactivity control system
2.1.3.1.4 Reactor pressure vessel
2.1.3.2 Reactor cooling system
2.1.3.3 Engineered safety systems
2.1.3.3.1 Auxiliary cooling system
2.1.3.3.2 Vessel cooling system
2.1.3.3.3 Containment structure
2.1.3.4 Instrumentation and control system
2.1.3.4.1 Instrumentation system
2.1.3.4.2 Control system
2.1.3.4.3 Safety protection system
2.1.4 R&D programs for HTTR
2.1.4.1 Fuel
2.1.4.2 Graphite
2.1.4.3 Metallic materials
2.1.4.4 Reactor physics
2.1.4.5 Reactor instrumentation
2.1.4.6 Heat transfer and fluid dynamics
2.1.4.6.1 Air ingress process following primary-pipe rupture
2.1.4.6.2 Graphite oxidation in case of air ingress into reactor core
2.1.4.7 Components and structures at high temperature
2.2 Nuclear design
2.2.1 Introduction
2.2.2 Design requirement
2.2.2.1 Excess reactivity
2.2.2.2 Reactor shutdown margin
2.2.2.3 Reactivity addition rate
2.2.2.4 Reactivity coefficient
2.2.2.5 Power distribution
2.2.2.6 Burnup
2.2.3 Analytical method
2.2.3.1 Design codes
2.2.3.2 Validation of design code using very high temperature reactor critical assembly
2.2.4 Evaluation of nuclear characteristics
2.2.4.1 Excess reactivity and nuclear shutdown margin
2.2.4.1.1 Excess reactivity
Reactivity losses
Excess reactivity
2.2.4.1.2 Burnable poison rods
2.2.4.1.3 Controllable reactivity and shutdown margin
Control rod
Reserved shutdown system
2.2.4.2 Reactivity addition rate and reactivity coefficient
2.2.4.2.1 Reactivity addition rate
2.2.4.2.2 Reactivity coefficient
2.2.4.3 Power distribution and burnup
2.2.4.3.1 Power distribution
Radial power distribution
Axial power distribution
2.2.4.3.2 Burnup
2.3 Core thermal-hydraulics
2.3.1 Introduction
2.3.2 Design requirements
2.3.3 Design details
2.3.4 Evaluation results of design
2.3.5 Reevaluation of maximum fuel temperature with operational data
2.3.5.1 Revision of calculation condition
2.3.5.2 Reevaluation result by operational data
2.4 Graphite components
2.4.1 Introduction
2.4.2 In-core graphite and carbon structure in high temperature engineering test reactor
2.4.2.1 Core graphite components
2.4.2.2 Core support graphite components
2.4.3 Concepts of graphite design criteria
2.4.3.1 Component classification
2.4.3.2 Fracture theory
2.4.3.3 Stress classification
2.4.3.4 Stress limit
2.4.3.5 Buckling limit
2.4.3.6 Stress analysis
2.4.3.7 Specified minimum ultimate strength
2.4.3.8 Oxidation effect
2.4.4 Quality control
2.5 Metallic components
2.5.1 Introduction
2.5.2 Development of Hastelloy XR
2.5.3 Identification of failure modes
2.5.4 Developments of design limits and rules
2.5.4.1 Hastelloy XR
2.5.4.1.1 Material characterization
2.5.4.1.2 Tensile property
2.5.4.1.3 Creep property
2.5.4.1.4 Creep–fatigue interaction
2.5.4.1.5 Applicability of the fast breeder reactor code
2.5.4.1.6 Structural mechanics behavior
2.5.4.1.7 Multiaxiality of creep rupture strength and creep–fatigue damage
2.5.4.1.8 Creep buckling
2.5.4.1.9 Creep analysis method
2.5.4.2 2¼ Cr–1Mo steel
2.5.4.3 Austenitic stainless steels SUS321TB and SUS316
2.5.4.4 1Cr–0.5Mo–V steel
2.6 Core components and reactor internals
2.6.1 Introduction
2.6.2 Fuel
2.6.2.1 Design requirement
2.6.2.2 Design details
2.6.2.3 Evaluation
2.6.3 Hexagonal graphite blocks
2.6.3.1 Design requirement
2.6.3.2 Design details
2.6.3.3 Evaluation
2.6.4 Core support structures
2.6.4.1 Design requirement
2.6.4.2 Design details
2.6.4.3 In-service inspection and surveillance test
2.6.4.3.1 In-service inspection using TV camera
2.6.4.3.2 Results of preservice inspection
2.6.4.3.3 Surveillance test
2.6.5 Core support metallic structures
2.6.5.1 Design requirement
2.6.5.2 Design details
2.6.5.3 In-service inspection and surveillance test
2.6.6 Shielding blocks
2.6.6.1 Design requirement
2.6.6.2 Design details
2.7 Seismic design
2.7.1 Introduction
2.7.2 Seismic design
2.7.2.1 Basic guideline of seismic design
2.7.2.2 Seismic classification
2.7.2.3 Basic design earthquake ground motion
2.7.3 Geological composition and seismometry
2.7.3.1 Geological composition
2.7.3.2 Seismometry
2.7.4 Structure of core components
2.7.5 Development of evaluation method
2.7.5.1 Vibration characteristics of core components
2.7.5.2 Validation of SONATINA-2V code
2.7.6 Structural integrity of graphite components
2.7.6.1 Core components
2.7.6.2 Core bottom structure
2.8 Cooling system
2.8.1 Introduction
2.8.2 Primary cooling system
2.8.2.1 Primary pressurized water cooler
2.8.2.2 Intermediate heat exchanger
2.8.2.3 Primary gas circulator
2.8.2.4 Primary concentric hot gas duct
2.8.3 Secondary helium cooling system
2.8.3.1 Secondary pressurized water cooler
2.8.3.2 Secondary gas circulator
2.8.3.3 Secondary helium piping
2.8.4 Pressurized water-cooling system
2.8.4.1 Pressurized water pump
2.8.4.2 Air cooler
2.8.5 Residual heat removal system
2.8.5.1 Auxiliary cooling system
2.8.5.2 Vessel cooling system
2.9 Reactivity control system
2.9.1 Introduction
2.9.2 Control rod system
2.9.2.1 Design requirement
2.9.2.2 Design details
2.9.2.2.1 Control rod
2.9.2.2.2 Control rod drive mechanism
2.9.2.3 High temperature structural design guideline of control rod
2.9.2.4 Design material data on Alloy 800H
2.9.2.5 Results of R&D
2.9.2.5.1 Scram tests of the control rod system under seismic conditions
2.9.2.5.2 Reliability test of control rods in the HENDEL loop
2.9.2.5.3 Verification tests of the control rods
2.9.2.6 Temperature analysis
2.9.2.7 Stress analysis
2.9.3 Reserve shutdown system
2.9.3.1 Design
2.9.3.2 Results of R&D
2.10 Instrumentation and control system
2.10.1 Introduction
2.10.2 Instrumentation system
2.10.2.1 Reactor instrumentation
2.10.2.1.1 Nuclear instrumentation
2.10.2.1.2 Control rods position instrumentation
2.10.2.1.3 Three core differential pressure instrumentation
2.10.2.1.4 Fuel failure detection system
2.10.2.1.5 In-core temperature monitoring system
2.10.3 Process instrumentation
2.10.4 Control system
2.10.4.1 Operational mode selector
2.10.4.2 Reactor power control device
2.10.4.2.1 Reactor power control system
2.10.4.2.2 Reactor outlet coolant temperature control system
2.10.4.3 Plant control device
2.10.4.3.1 Reactor inlet coolant temperature control system
2.10.4.3.2 Intermediate heat exchanger primary coolant flow rate control system
2.10.4.3.3 Primary-pressurized water cooler primary coolant flow rate control system
2.10.4.3.4 Primary helium pressure control system
2.10.4.3.5 Primary–secondary helium differential pressure control system
2.10.4.3.6 Primary-pressurized water differential pressure control system
2.10.4.3.7 Pressurized water temperature control system
2.10.5 Safety protection system
2.10.5.1 Reactor protection system
2.10.5.2 Engineered safety features actuating system
2.10.5.2.1 Signal isolating containment vessel
2.10.5.2.2 Signal starting up auxiliary cooling system
2.10.5.2.3 Signal isolating auxiliary cooling water line
2.10.6 Performance test results
2.10.6.1 Characteristics of the neutron flux monitoring system
2.10.6.2 Primary-pressurized water cooler primary coolant flow rate control system
2.10.6.3 Reactor outlet coolant temperature control system
2.11 Containment structures
2.11.1 Introduction
2.11.2 Reactor containment vessel
2.11.2.1 Design and construction
2.11.2.2 Leakage-rate test
2.11.2.2.1 Partial leakage-rate test
2.11.2.2.2 Whole leakage-rate test
2.11.3 Service area
2.11.3.1 Design
2.11.3.2 Commissioning tests
2.11.4 Emergency air purification system
2.11.4.1 Design
2.11.4.2 Commissioning tests
2.11.4.2.1 Start-up test
2.11.4.2.2 Filter efficiency measurement
2.12 Other systems
2.12.1 Introduction
2.12.2 Auxiliary helium systems
2.12.2.1 Helium purification system
2.12.2.2 Helium sampling system
2.12.2.3 Helium storage and supply system
2.12.3 Fuel system
2.12.3.1 Fuel handling system
2.12.3.2 Fuel storage system
2.13 Safety design
2.13.1 Introduction
2.13.2 Basic safety design philosophy
2.13.3 Safety classification
2.13.4 Fundamental safety functions unique to HTTR
2.13.4.1 Control of reactivity
2.13.4.2 Removal of heat from core
2.13.4.3 Confinement of fission product release
2.13.5 Acceptance criteria
2.13.6 Selection of events
2.13.7 Safety evaluation technologies
2.13.8 New safety criteria
References
3---R-amp-D-on-components_2021_High-Temperature-Gas-Cooled-Reactors
3 R&D on components
3.1 Fuel
3.1.1 Introduction
3.1.2 Related research and development for fuel design
3.1.2.1 Limitation for as-fabricated fuel failure fraction
3.1.2.2 Kernel migration
3.1.2.3 Palladium (Pd)–SiC interaction
3.1.2.4 Confirmation up to maximum burnup
3.1.2.5 Temperature limit of HTTR fuel
3.1.3 Fabrication technologies for HTTR fuel
3.1.3.1 Improvement of fuel fabrication process
3.1.3.2 Quality control
3.1.4 Performance of HTTR fuel during long-term high temperature operation
3.2 Core components and reactor internals
3.2.1 Introduction
3.2.2 Tests on core components
3.2.2.1 Test apparatus
3.2.2.2 Thermal hydraulic characteristics
3.2.2.2.1 Pressure drop
3.2.2.2.2 Heat transfer
3.2.2.3 Effect of power unbalance
3.2.3 Tests on reactor internals
3.2.3.1 Test apparatus
3.2.3.2 Sealing performance of helium gas
3.2.3.3 Mixing performance of helium gas
3.2.3.4 Insulation performance of core bottom structure
3.2.3.5 Thermal performance of a coaxial hot gas duct
3.3 Passive cooling system
3.3.1 Introduction
3.3.2 Experiment
3.3.2.1 Experimental apparatus
3.3.2.2 Measurement
3.3.2.3 Experimental conditions
3.3.3 Numerical method
3.3.3.1 Numerical code
3.3.3.2 Numerical model
3.3.3.3 Empirical correlations for natural convective heat transfer
3.3.4 Evaluation of hot spot by natural convection
3.3.5 Evaluation of local hot spot around standpipes
3.4 Intermediate heat exchanger
3.4.1 Introduction
3.4.2 Creep collapse of the tube against external pressure
3.4.2.1 Objective and test procedure
3.4.2.2 Test results
3.4.3 Creep fatigue of tube against thermal stress
3.4.3.1 Objective and test procedure
3.4.3.2 Test results
3.4.4 Seismic behavior of tube bundle
3.4.4.1 Objective and test procedure
3.4.4.2 Test results
3.4.5 Thermal hydraulic behavior of tube bundle
3.4.5.1 Objective and test procedure
3.4.5.2 Test results
3.4.6 In-service inspection technology of tube
3.4.6.1 Objective and test procedure
3.4.6.2 Test results
3.5 Basic feature of air ingress during primary pipe rupture accident
3.5.1 Introduction
3.5.2 Basic feature of air ingress phenomena in a reverse U-shaped channel
3.5.2.1 Experimental apparatus, method, and results
3.5.2.2 Existence of parallel channels in a reactor core
3.5.2.2.1 Isothermal condition
3.5.2.2.2 Different temperature condition
3.5.3 Basic feature of air ingress phenomena in a simulated reactor apparatus
3.5.3.1 Numerical analysis
3.5.3.2 Comparison with experiments
References
4---Operation-of-HTTR_2021_High-Temperature-Gas-Cooled-Reactors
4 Operation of HTTR
4.1 Unexpected incidents under construction and operation
4.1.1 Introduction
4.1.2 Temperature rise of primary upper shielding
4.1.2.1 Outline of incident
4.1.2.2 First countermeasures to reduce temperature
4.1.2.2.1 Countermeasure
4.1.2.2.2 Test result after first countermeasure
4.1.2.3 Second countermeasures to reduce temperature
4.1.2.3.1 Countermeasure
4.1.2.3.2 Test result after installation of second countermeasure
4.1.2.4 Prediction and test results at full power
4.1.3 Temperature rise of core support plate
4.1.3.1 Outline of incident
4.1.3.2 Reevaluation of core support plate temperature
4.2 Characteristic test of initial core
4.2.1 Introduction
4.2.2 General description
4.2.2.1 Test condition
4.2.2.2 Nuclear calculations
4.2.3 Critical approach
4.2.4 Excess reactivity and shutdown margin
4.2.4.1 Excess reactivity
4.2.4.2 Shutdown margin
4.2.5 Control rod characteristics
4.2.5.1 Control rod worth on zero power condition
4.2.5.2 Control rod position versus reactor power
4.2.6 Reactivity coefficient
4.2.6.1 Temperature coefficient
4.2.6.2 Power coefficient
4.2.7 Neutron flux and power distribution
4.2.7.1 Neutron flux distribution
4.2.7.2 Power distribution
4.3 Performance test
4.3.1 Introduction
4.3.2 Major test items
4.3.3 Heat balance of reactor cooling system
4.3.4 Heat exchanger performance
4.3.5 Reactor control system performance
4.3.6 Residual heat removal performance at manual reactor scram
4.3.7 Thermal expansion performance of high temperature components
4.3.8 Fuel and fission product behavior
4.4 High temperature operation
4.4.1 Introduction
4.4.2 Main test results of long-term high temperature operation
4.4.2.1 30-Day continuous operation
4.4.2.2 50-Day continuous operation
4.4.3 Validation using high temperature engineering test reactor burnup data
4.4.3.1 Trend of change in control rod position
4.4.3.2 Effectiveness of rod-type burnable poisons
4.4.3.2.1 Design philosophy of burnable poisons
4.4.3.2.2 Validity of effectiveness of rod-type burnable poisons
4.4.3.3 Whole core burnup calculations
4.4.3.3.1 Calculation method
4.4.3.3.2 Validity of whole core burnup calculations
4.5 Safety demonstration test
4.5.1 Introduction
4.5.2 High temperature engineering test reactor control system
4.5.2.1 Reactivity and reactor power control systems
4.5.2.2 Cooling system and plant control device
4.5.3 Safety demonstration test plan
4.5.4 Analysis code and model
4.5.5 Reactivity insertion test
4.5.5.1 Objective and test procedure
4.5.5.2 Test results
4.5.6 Coolant flow reduction test—gas circulators trip test
4.5.6.1 Objective and test procedure
4.5.6.2 Test results
4.5.7 Loss of forced cooling test
4.5.7.1 Objective and test procedure
4.5.7.2 Test results
References
5---R-amp-D-on-commercial-high-temperature-ga_2021_High-Temperature-Gas-Cool
5 R&D on commercial high temperature gas-cooled reactor
5.1 System design for power generation
5.1.1 Introduction
5.1.2 HTR50S: HTGR steam cycle power plant
5.1.3 GTHTR300: HTGR gas turbine power plant
5.2 System design for cogeneration
5.2.1 Introduction
5.2.2 Hydrogen cogeneration
5.2.3 Seawater desalination
5.2.4 HTGR renewable hybrid system
5.3 System design for steelmaking
5.3.1 Introduction
5.3.2 Flow diagram of steelmaking systems
5.3.3 CO2 emission
5.3.4 Steelmaking cost
5.4 Safety design for connection of heat application system and high temperature gas-cooled reactor
5.4.1 Introduction
5.4.2 Roadmap for safety standard establishment
5.4.3 Safety requirements
5.4.3.1 Basic safety approach
5.4.3.2 Safety requirements
5.4.3.2.1 Confinement of radionuclides
5.4.3.2.2 Control reactivity
5.4.3.2.3 Heat removal from core
5.4.3.2.4 Loss-of-offsite power
5.4.3.2.5 Coupling to heat application system
5.4.4 Basic concept of safety guides
5.4.4.1 Evaluation items
5.4.4.2 Licensing basis event selection
5.4.4.2.1 Identification of abnormal events and postulated initiating events
5.4.4.2.2 Definition of safety function and mitigation system
5.4.4.2.3 Grouping of abnormal events
5.4.4.2.4 Licensing basis event selection for single failure events
5.4.4.2.5 Identification of accident sequence
5.4.4.2.6 Grouping of accident sequence
5.4.4.2.7 Identification of significant accident sequence
5.4.4.2.8 Licensing basis event selection
5.4.4.3 Acceptance criteria
5.4.4.3.1 Anticipated operational occurrences
5.4.4.3.2 Accidents
5.4.5 HTTR cogeneration demonstration
5.5 Gas turbine technology for power generation
5.5.1 Introduction
5.5.2 Helium gas compressor
5.5.3 Magnetic bearing
5.6 Iodine–sulfur process technology for hydrogen production
5.6.1 Introduction
5.6.2 Bench-scale test
5.6.3 Elemental technologies
5.6.4 Industrial material component test
5.6.5 Hydrogen production test
5.6.6 Improvement of hydrogen production efficiency
5.6.6.1 Electro-electrodialysis cell
5.6.6.2 Hydrogen iodide decomposer with hydrogen separation membrane
5.6.6.3 Analytical estimation of hydrogen production thermal efficiency
5.6.7 Component materials
5.6.7.1 Corrosion resistance
5.6.7.2 Strength
5.7 System integration technology for connection of heat application system and high temperature gas-cooled reactor
5.7.1 Introduction
5.7.2 Control technology
5.7.3 Tritium permeation
5.7.4 Explosion of combustible gas
5.7.5 High temperature isolation valves
5.8 Prevention technology for air ingress during a primary pipe rupture accident
5.8.1 Introduction
5.8.2 Prevention technology of air ingress in a reverse U-shaped channel
5.8.2.1 Experimental apparatus, method, and results
5.8.2.2 Numerical analysis
5.8.2.3 Onset time of natural circulation flow through the apparatus
5.8.2.4 Onset time of natural circulation
5.8.3 Basic feature of air ingress phenomena during a horizontal pipe break accident
5.8.3.1 Introduction
5.8.3.2 Experimental apparatus
5.8.3.3 Experimental method
5.8.3.4 Experimental results
5.9 Advanced fuel technology for high burnup
5.9.1 Introduction
5.9.2 Design of high burnup fuel
5.9.3 Upgrade technologies for high burnup
5.9.3.1 Fuel design
5.9.3.2 Irradiation test
5.9.4 Future study plan
5.10 Advanced fuel for plutonium burner
5.10.1 Introduction
5.10.2 Fuel fabrication process of Clean Burn
5.10.3 Core design
5.10.4 Future study plan
5.11 Advanced fuel technology for reduction of high-level radioactive waste
5.11.1 Introduction
5.11.2 Calculation for repository design
5.11.3 Evaluation of waste package
References
Index_2021_High-Temperature-Gas-Cooled-Reactors
Index

Citation preview

High Temperature Gas-cooled Reactors

JSME Series in Thermal and Nuclear Power Generation

High Temperature Gas-cooled Reactors Edited by

Tetsuaki Takeda Graduate Faculty of Interdisciplinary Research, Research Faculty of Engineering, Department of Mechanical Engineering, University of Yamanashi, Kofu, Japan

Yoshiyuki Inagaki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Oarai, Japan

Series Editor

Yasuo Koizumi Graduate School of Information and Engineering, The University of Electro-Communications, Tokyo, Japan

Elsevier Radarweg 29, PO Box 211, 1000 AE Amsterdam, Netherlands The Boulevard, Langford Lane, Kidlington, Oxford OX5 1GB, United Kingdom 50 Hampshire Street, 5th Floor, Cambridge, MA 02139, United States Copyright © 2021 Elsevier Inc. All rights reserved. No part of this publication may be reproduced or transmitted in any form or by any means, electronic or mechanical, including photocopying, recording, or any information storage and retrieval system, without permission in writing from the publisher. Details on how to seek permission, further information about the Publisher’s permissions policies and our arrangements with organizations such as the Copyright Clearance Center and the Copyright Licensing Agency, can be found at our website: www.elsevier.com/permissions. This book and the individual contributions contained in it are protected under copyright by the Publisher (other than as may be noted herein). Notices Knowledge and best practice in this field are constantly changing. As new research and experience broaden our understanding, changes in research methods, professional practices, or medical treatment may become necessary. Practitioners and researchers must always rely on their own experience and knowledge in evaluating and using any information, methods, compounds, or experiments described herein. In using such information or methods they should be mindful of their own safety and the safety of others, including parties for whom they have a professional responsibility. To the fullest extent of the law, neither the Publisher nor the authors, contributors, or editors, assume any liability for any injury and/or damage to persons or property as a matter of products liability, negligence or otherwise, or from any use or operation of any methods, products, instructions, or ideas contained in the material herein. British Library Cataloguing-in-Publication Data A catalogue record for this book is available from the British Library Library of Congress Cataloging-in-Publication Data A catalog record for this book is available from the Library of Congress ISBN: 978-0-12-821031-4 For Information on all Elsevier publications visit our website at https://www.elsevier.com/books-and-journals

Publisher: Brian Romer Acquisitions Editor: Maria Convey Editorial Project Manager: Sara Valentino Production Project Manager: Kamesh Ramajogi Cover Designer: Alan Studholme Typeset by MPS Limited, Chennai, India

Contents

List of contributors About the authors Preface of JSME Series in Thermal and Nuclear Power Generation Preface to Volume 5: High-Temperature Gas-Cooled Reactors 1.

2.

xi xv xxix xxxiii

Overview of high temperature gas-cooled reactor Jin Iwatsuki, Kazuhiko Kunitomi, Hideaki Mineo, Tetsuo Nishihara, Nariaki Sakaba, Masayuki Shinozaki, Yukio Tachibana and Xing Yan 1.1 Features of high temperature gas-cooled reactor 1.1.1 Structure and materials 1.1.2 Heat application 1.1.3 Safety 1.1.4 Adaptability to environment 1.2 History of research and development in world 1.3 History of research and development in Japan References Design of High Temperature Engineering Test Reactor (HTTR) Yusuke Fujiwara, Minoru Goto, Kazuhiko Iigaki, Tatsuo Iyoku, Hai Quan Ho, Taiki Kawamoto, Makoto Kondo, Kazuhiko Kunitomi, Keisuke Morita, Satoru Nagasumi, Shigeaki Nakagawa, Tetsuo Nishihara, Naoki Nojiri, Masato Ono, Akio Saikusa, Nariaki Sakaba, Taiju Shibata, Yosuke Shimazaki, Atsushi Shimizu, Masayuki Shinozaki, Junya Sumita, Yukio Tachibana, Shoji Takada, Daisuke Tochio, Takahiro Uesaka and Shohei Ueta 2.1 Overview of HTTR design features 2.1.1 Introduction 2.1.2 History and future plan of HTTR project 2.1.3 Major design features of HTTR 2.1.4 R&D programs for HTTR 2.2 Nuclear design 2.2.1 Introduction 2.2.2 Design requirement 2.2.3 Analytical method 2.2.4 Evaluation of nuclear characteristics 2.3 Core thermal-hydraulics 2.3.1 Introduction

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Contents

2.3.2 2.3.3 2.3.4 2.3.5

2.4

2.5

2.6

2.7

2.8

2.9

2.10

Design requirements Design details Evaluation results of design Reevaluation of maximum fuel temperature with operational data Graphite components 2.4.1 Introduction 2.4.2 In-core graphite and carbon structure in high temperature engineering test reactor 2.4.3 Concepts of graphite design criteria 2.4.4 Quality control Metallic components 2.5.1 Introduction 2.5.2 Development of Hastelloy XR 2.5.3 Identification of failure modes 2.5.4 Developments of design limits and rules Core components and reactor internals 2.6.1 Introduction 2.6.2 Fuel 2.6.3 Hexagonal graphite blocks 2.6.4 Core support structures 2.6.5 Core support metallic structures 2.6.6 Shielding blocks Seismic design 2.7.1 Introduction 2.7.2 Seismic design 2.7.3 Geological composition and seismometry 2.7.4 Structure of core components 2.7.5 Development of evaluation method 2.7.6 Structural integrity of graphite components Cooling system 2.8.1 Introduction 2.8.2 Primary cooling system 2.8.3 Secondary helium cooling system 2.8.4 Pressurized water-cooling system 2.8.5 Residual heat removal system Reactivity control system 2.9.1 Introduction 2.9.2 Control rod system 2.9.3 Reserve shutdown system Instrumentation and control system 2.10.1 Introduction 2.10.2 Instrumentation system 2.10.3 Process instrumentation 2.10.4 Control system

41 42 44 45 48 48 48 50 58 58 58 59 61 61 71 71 71 74 79 85 87 89 89 90 94 94 99 100 102 102 103 110 112 113 113 113 114 126 127 127 128 131 132

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2.10.5 Safety protection system 2.10.6 Performance test results 2.11 Containment structures 2.11.1 Introduction 2.11.2 Reactor containment vessel 2.11.3 Service area 2.11.4 Emergency air purification system 2.12 Other systems 2.12.1 Introduction 2.12.2 Auxiliary helium systems 2.12.3 Fuel system 2.13 Safety design 2.13.1 Introduction 2.13.2 Basic safety design philosophy 2.13.3 Safety classification 2.13.4 Fundamental safety functions unique to HTTR 2.13.5 Acceptance criteria 2.13.6 Selection of events 2.13.7 Safety evaluation technologies 2.13.8 New safety criteria References

134 135 138 138 138 146 147 151 151 151 155 158 158 158 160 164 165 167 169 173 173

R&D on components Jun Aihara, Minoru Goto, Yoshiyuki Inagaki, Tatsuo Iyoku, Kazuhiko Kunitomi, Tetsuo Nishihara, Nariaki Sakaba, Taiju Shibata, Junya Sumita, Yukio Tachibana, Shoji Takada, Tetsuaki Takeda and Shohei Ueta 3.1 Fuel 3.1.1 Introduction 3.1.2 Related research and development for fuel design 3.1.3 Fabrication technologies for HTTR fuel 3.1.4 Performance of HTTR fuel during long-term high temperature operation 3.2 Core components and reactor internals 3.2.1 Introduction 3.2.2 Tests on core components 3.2.3 Tests on reactor internals 3.3 Passive cooling system 3.3.1 Introduction 3.3.2 Experiment 3.3.3 Numerical method 3.3.4 Evaluation of hot spot by natural convection 3.3.5 Evaluation of local hot spot around standpipes 3.4 Intermediate heat exchanger 3.4.1 Introduction 3.4.2 Creep collapse of the tube against external pressure

179

180 180 181 185 189 191 191 191 196 203 203 205 208 213 215 217 217 218

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3.4.3 Creep fatigue of tube against thermal stress 3.4.4 Seismic behavior of tube bundle 3.4.5 Thermal hydraulic behavior of tube bundle 3.4.6 In-service inspection technology of tube 3.5 Basic feature of air ingress during primary pipe rupture accident 3.5.1 Introduction 3.5.2 Basic feature of air ingress phenomena in a reverse U-shaped channel 3.5.3 Basic feature of air ingress phenomena in a simulated reactor apparatus References

220 223 226 228 230 230

Operation of HTTR Yuji Fukaya, Minoru Goto, Hiroyuki Inoi, Etsuo Ishitsuka, Tatsuo Iyoku, Kazuhiko Kunitomi, Shigeaki Nakagawa, Tetsuo Nishihara, Hai Quan Ho, Akio Saikusa, Nariaki Sakaba, Hiroaki Sawahata, Taiju Shibata, Masayuki Shinozaki, Yukio Tachibana, Shoji Takada, Kuniyoshi Takamatsu and Daisuke Tochio 4.1 Unexpected incidents under construction and operation 4.1.1 Introduction 4.1.2 Temperature rise of primary upper shielding 4.1.3 Temperature rise of core support plate 4.2 Characteristic test of initial core 4.2.1 Introduction 4.2.2 General description 4.2.3 Critical approach 4.2.4 Excess reactivity and shutdown margin 4.2.5 Control rod characteristics 4.2.6 Reactivity coefficient 4.2.7 Neutron flux and power distribution 4.3 Performance test 4.3.1 Introduction 4.3.2 Major test items 4.3.3 Heat balance of reactor cooling system 4.3.4 Heat exchanger performance 4.3.5 Reactor control system performance 4.3.6 Residual heat removal performance at manual reactor scram 4.3.7 Thermal expansion performance of high temperature components 4.3.8 Fuel and fission product behavior 4.4 High temperature operation 4.4.1 Introduction 4.4.2 Main test results of long-term high temperature operation 4.4.3 Validation using high temperature engineering test reactor burnup data

257

233 243 253

258 258 258 264 269 269 270 271 272 274 275 276 278 278 279 280 281 282 284 285 286 287 287 288 292

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5.

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4.5

Safety demonstration test 4.5.1 Introduction 4.5.2 High temperature engineering test reactor control system 4.5.3 Safety demonstration test plan 4.5.4 Analysis code and model 4.5.5 Reactivity insertion test 4.5.6 Coolant flow reduction test—gas circulators trip test 4.5.7 Loss of forced cooling test References

298 298 300 302 303 304 307 309 310

R&D on commercial high temperature gas-cooled reactor Jun Aihara, Takeshi Aoki, Yuji Fukaya, Minoru Goto, Yoshiyuki Imai, Yoshitomo Inaba, Yoshiyuki Inagaki, Tatsuo Iyoku, Yu Kamiji, Seiji Kasahara, Shinji Kubo, Kazuhiko Kunitomi, Naoki Mizuta, Odtsetseg Myagmarjav, Tetsuo Nishihara, Hiroki Noguchi, Hirofumi Ohashi, Nariaki Sakaba, Koei Sasaki, Hiroyuki Sato, Taiju Shibata, Junya Sumita, Yukio Tachibana, Shoji Takada, Tetsuaki Takeda, Hiroaki Takegami, Nobuyuki Tanaka, Shohei Ueta and Xing Yan 5.1 System design for power generation 5.1.1 Introduction 5.1.2 HTR50S: HTGR steam cycle power plant 5.1.3 GTHTR300: HTGR gas turbine power plant 5.2 System design for cogeneration 5.2.1 Introduction 5.2.2 Hydrogen cogeneration 5.2.3 Seawater desalination 5.2.4 HTGR renewable hybrid system 5.3 System design for steelmaking 5.3.1 Introduction 5.3.2 Flow diagram of steelmaking systems 5.3.3 CO2 emission 5.3.4 Steelmaking cost 5.4 Safety design for connection of heat application system and high temperature gas-cooled reactor 5.4.1 Introduction 5.4.2 Roadmap for safety standard establishment 5.4.3 Safety requirements 5.4.4 Basic concept of safety guides 5.4.5 HTTR cogeneration demonstration 5.5 Gas turbine technology for power generation 5.5.1 Introduction 5.5.2 Helium gas compressor 5.5.3 Magnetic bearing

313

314 314 315 319 329 329 330 336 338 341 341 341 346 346 349 349 349 351 353 355 359 359 360 366

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Contents

5.6

Iodine sulfur process technology for hydrogen production 5.6.1 Introduction 5.6.2 Bench-scale test 5.6.3 Elemental technologies 5.6.4 Industrial material component test 5.6.5 Hydrogen production test 5.6.6 Improvement of hydrogen production efficiency 5.6.7 Component materials 5.7 System integration technology for connection of heat application system and high temperature gas-cooled reactor 5.7.1 Introduction 5.7.2 Control technology 5.7.3 Tritium permeation 5.7.4 Explosion of combustible gas 5.7.5 High temperature isolation valves 5.8 Prevention technology for air ingress during a primary pipe rupture accident 5.8.1 Introduction 5.8.2 Prevention technology of air ingress in a reverse U-shaped channel 5.8.3 Basic feature of air ingress phenomena during a horizontal pipe break accident 5.9 Advanced fuel technology for high burnup 5.9.1 Introduction 5.9.2 Design of high burnup fuel 5.9.3 Upgrade technologies for high burnup 5.9.4 Future study plan 5.10 Advanced fuel for plutonium burner 5.10.1 Introduction 5.10.2 Fuel fabrication process of Clean Burn 5.10.3 Core design 5.10.4 Future study plan 5.11 Advanced fuel technology for reduction of high-level radioactive waste 5.11.1 Introduction 5.11.2 Calculation for repository design 5.11.3 Evaluation of waste package References Index

370 370 371 371 372 377 379 383 387 387 389 393 396 397 399 399 399 415 421 421 422 424 429 429 429 430 432 436 437 437 437 441 445 451

List of contributors

Jun Aihara Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Takeshi Aoki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yusuke Fujiwara Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yuji Fukaya Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Minoru Goto Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Hai Quan Ho Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Kazuhiko Iigaki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yoshiyuki Imai Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yoshitomo Inaba Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yoshiyuki Inagaki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Hiroyuki Inoi Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Etsuo Ishitsuka Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

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List of contributors

Jin Iwatsuki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Tatsuo Iyoku Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yu Kamiji Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Seiji Kasahara Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Taiki Kawamoto Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Makoto Kondo Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Shinji Kubo Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Kazuhiko Kunitomi Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Hideaki Mineo Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Naoki Mizuta Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Keisuke Morita Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Odtsetseg Myagmarjav Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Satoru Nagasumi Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Shigeaki Nakagawa Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Tetsuo Nishihara Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

List of contributors

xiii

Hiroki Noguchi Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Naoki Nojiri Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Hirofumi Ohashi Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Masato Ono Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Akio Saikusa Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Nariaki Sakaba Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Koei Sasaki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Hiroyuki Sato Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Hiroaki Sawahata Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Taiju Shibata Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yosuke Shimazaki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Atsushi Shimizu Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Masayuki Shinozaki Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Junya Sumita Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Yukio Tachibana Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

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List of contributors

Shoji Takada Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Kuniyoshi Takamatsu Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Tetsuaki Takeda Graduate Faculty of Interdisciplinary Research, Research Faculty of Engineering, Department of Mechanical Engineering, University of Yamanashi, Yamanashi, Japan Hiroaki Takegami Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Nobuyuki Tanaka Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Daisuke Tochio Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Takahiro Uesaka Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Shohei Ueta Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan Xing Yan Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

About the authors

Jun Aihara is Assistant Principal Researcher of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. She received her PhD degree from Ibaraki University in 2008. She started her research career at Japan Atomic Energy Research Institute (JAERI) in 1996. Her research focuses on the areas of fuel of high-temperature gas-cooled reactors. She is also interested in microstructural observation. Takeshi Aoki is Research Engineer of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Tokyo Institute of Technology in 2018. He has been in his present position from 2018. He works on the thermal hydraulic and safety design and analysis for hightemperature gas-cooled reactor. His current works mainly focus on studies on thermal hydraulic behaviors in normal operation and accident condition. His research interests also include nuclear security and nonproliferation features of the nuclear reactor system. Yusuke Fujiwara is Research Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Osaka University in 2014. He started his research career at JAEA in 2014 for the research and development of the HTGR technology. He is mainly interested in mechanical engineering of HTGR. His current task is operation and maintenance of the HTTR, including an evaluation of external event, the effect of Tornado on the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. Yuji Fukaya is Principal Researcher of Reactor Systems Design Department, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from The University of Tokyo in 2009. He started his research career at JAERI in 2002 as a research engineer for nuclear criticality safety. He then starts the core design of reduced moderation water reactor, fast breeder reactor (FBR), and HTGR, respectively, in 2005, 2010, and 2011. His research focuses on the wide area not only of core physics but also of sustainability

xvi

About the authors

of frontend, environmental burden of backend, nonproliferation, and economy of the fuel cycle, since his research field is closely related to the fuel cycle system. Minoru Goto is Group Leader of Reactor Systems Design Department, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Kyushu University in 2014. He worked in a private company until 2002 mainly on the development of analysis codes for reprocessing plants and radiation shielding analyses. In 2003 he joined JAERI and started his research career for the nuclear reactor design of HTGR. His research focuses on the areas of the nuclear and thermal design of HTGR and applications using HTGR. He is also interested in nuclear data and reprocessing. Hai Quan Ho is Research Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Tokyo Institute of Technology in 2016. His current research focuses on the neutronic and burn up calculation of HTGR. His interest is also to expand the multipurpose use of HTGR. He investigates the feasibility of some neutron irradiation applications using HTGR, such as silicon neutron transmutation doping or radioisotopes production. Kazuhiko Iigaki is Manager of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Tokai University in 1995. He started his research career to join the HTTR project at JAERI in 1995. He has experienced the construction, operation, and maintenance of the HTTR. He joined Nuclear Safety Commission from 2006 to 2008 to review the license of nuclear facilities. Since 2008 he joined the HTTR project to carry out a seismic back-check of the HTTR, and a seismic evaluation of the HTTR reactor facility as well as BDBA for the conformity review by Nuclear Regulation Authority on the new regulatory requirements that were issued in 2014. Yoshiyuki Imai is Research Engineer of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Kyoto University in 2005. He started his research career at JAERI in 2005 as a research engineer for hydrogen production. His research focuses on the areas of chemical engineering. He is also interested in materials to reduce fission product deposition. Since his research field is closely related to energy economics, he has great interest in heat application and hydrogen supply. Yoshitomo Inaba is Principal Researcher of International Cooperation and Social Environmental Office, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Kyushu University in

About the authors

xvii

2007. He started his research career at JAERI in 1995 for thermal and hydraulics on HTGR. He then worked at the Forschungszentrum Karlsruhe in Germany from 2004 to 2005. He has been in his present position since 2018. His work focuses on the areas of international cooperation on HTGR technologies. Yoshiyuki Inagaki is Researcher of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Kyushu University in 1996. He started his research career at JAERI in 1981 for the HTGR technology such as reactor internals and high-temperature components. He then worked in Juelich Research Center in Germany from 1994 to 1995 for a hydrogen production process. He also researched HTGR heat-application systems such as hydrogen production and nuclear steelmaking systems and the system integration technology for safe connection between a reactor and a hydrogen production system. Hiroyuki Inoi is Assistant Principal Researcher of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his BA degree from Kinki University in 2004. He started his research career to join the HTTR project for the research and development of the HTGR technology. He is engaged in the maintenance of BOP of the HTTR. His current task is to organize the conformity review of the HTTR by Nuclear Regulation Authority on the new regulatory requirements. He is also engaged in the operational safety supervision of the HTTR. Etsuo Ishitsuka is General Manager of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from The University of Tokyo in 1999. He started his research career at JAERI in 1986. He worked in a wide field of the development of neutron irradiation technology, such as the production of medical radioisotope and NTD silicon semiconductor, structural materials of light water reactor, fusion blanket materials, plasma-facing components, plasma diagnostics components, and measuring devices. He has managed the experiments of a fusion blanket functional test in the JMTR and the ITER project as the deputy general manager. After managing an international cooperation of the JMTR and training foreign young researchers/engineers using the JMTR, he joined the HTTR project in 2015. His current interest is the neutron irradiation technology of HTGR and its new applications. Jin Iwatsuki is Assistant Principal Researcher of Strategy and Planning Office in JAEA. He received his PhD degree from Tohoku University in 2012. He is responsible for the strategy of the HTGR project. His research focuses on the areas of design of chemical and nuclear heat utilization plants. He is also interested in

xviii

About the authors

design of chemical process equipment. Since his research field is closely related to chemical engineering, he has great interest in thermochemical hydrogen production process. Tatuso Iyoku is Associate Researcher of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from the University of Gunma in 1979. He experienced the design, construction, installation, inspection, testing, and in-service operation of the HTTR. His research mainly focuses on the mechanical design of graphite and graphite oxidation. He had been responsible for the maintenance and operation of the HTTR from April 2012 to 2014. Yu Kamiji is Research Engineer of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Utsunomiya University in 2010. He started his research career at JAEA in 2010. His research focuses on improving integrity for IS processes such as corrosion resistance of metallic materials, plant operation, and quality management of the components. He is also interested in hydrogen mitigation technology, especially the development of passive autocatalytic recombiner and the prediction of hydrogen advection and diffusion. Seiji Kasahara is Assistant Principal Researcher of Reactor Systems Design Department, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from The University of Tokyo in 2002. He started his research career at JAERI in 2002 in chemical engineering. His research focuses on the areas of design of chemical and steelmaking plants. He is also interested in design of chemical process equipment. Since his research field is closely related to chemical processes, he has great interest in thermochemical hydrogen production process. Taiki Kawamoto is Chief Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received BA degree from Tokai University in 2001. In the same year, he joined JAERI and was assigned to the HTTR as a reactor operator and maintenance staff. He joined the inspection of control rod, the replacement of neutron source and neutron detector, and the inspection of helium gas circulator and pressurized water cooler. His primary interest is the development of operation and maintenance technologies of the HTTR. Makoto Kondo is Chief Engineer of the Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast

About the authors

xix

Reactor and Advanced Reactor Research and Development in JAEA. He received his BA degree from Kinki University in 2002. He started his research career to join the HTTR project at JAERI in 2002 for the research and development of the HTGR technology. He is mainly interested in reactor operation and maintenance of HTGRs. He seconded the Nuclear Safety Commission of the Cabinet Office for researching nuclear regulation from 2009 to 2012. His current task is operation and maintenance of the HTTR, including an evaluation of external fire around the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. Shinji Kubo is Senior Principal Researcher of Department of Hydrogen and Heat Application Research and Development, Oarai Research and Development Institute, HTGR Research and Development Center, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Utsunomiya University in 1993. He started his research career at JAERI for heat utilization technologies using HTGR. His research focuses on the areas of thermochemical iodine sulfur hydrogen production process. He achieved world’s first stable hydrogen production for 1 week using glass equipment in 2004; subsequently, he developed three main reactors manufactured by industrial materials for use in high-temperature and highly corrosive environments and succeeded in hydrogen production operation for 150 h in 2019. Now, he leads the R&D program of hydrogen production process using HTGR in JAEA. Kazuhiko Kunitomi is Deputy Director General of Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD from Osaka University in 1990. He then worked in MIT from 1990 to 1991 for research on HTGR. After conducting research and development on HTGR internal core structures and primary components using HENDEL, he participated in the HTTR project and worked mainly on safety analysis, safety evaluation, and planning of the commissioning test for the HTTR. Then he moved to the future HTGR design and development section and designed several commercial HTGR systems with collaboration from Japanese industries. In addition, he was responsible for research and development program on gas turbine system for HTGR. His experiences extend to international research activities in IAEA, Generation IV International Forum (GIF), OECD, etc. His specialty is HTGR safety, thermal hydraulic analysis, and structural analysis. Now, he leads the HTGR development program in Japan. Hideaki Mineo is Director General of HTGR Research and Development Center and Deputy Director General of Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA from 2018. He received his PhD degree from The University of Tokyo in 1992. He started his research career at the Chemical Engineering Department of the university as a research assistant for fluidization engineering from 1988 to 1990. Also, he worked at the University of Bradford in United Kingdom during this period. In 1991 he joined JAEA. His research focused on the R&D of advanced spent fuel

xx

About the authors

reprocessing. He also worked in STA and Atomic Energy Commission of Japan. In 2014 he became a Director of Policy and Planning Department of JAEA. Since his research field is chemical engineering, he has great interest in the IS process and nuclear fuel cycle for HTGRs. Naoki Mizuta is Research Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Tokyo Institute of Technology in 2016. He started his research career at JAEA in 2016 for the HTGR fuel and materials. His current research is to develop an HTGR fuel compact with high oxidation resistance. He has large interest in neutron irradiation behavior and oxidation behavior of ceramic for hightemperature applications. Keisuke Morita is Research Engineer of Department of HTTR, HTGR Research and Development Center, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Kyushu University in 2018. He started his research career to join the HTTR project at JAERI in 2018 as a research engineer for the research and development of the HTGR technology. He is mainly interested in nuclear reactor physics of HTGRs. His current task is the operation and maintenance of the HTTR in the field of instrumentation and control system. Odtsetseg Myagmarjav is Research Engineer of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. She received her PhD degree from Tokyo Institute of Technology in 2015. She began her career as a postdoctoral fellow in JAEA where she specialized in nuclear hydrogen production via thermochemical processes, silica membrane synthesis, and membrane reactors. That eventually led to her position as a full researcher at the same institution in 2017. Her main research interests are the development of materials and devices for sustainable clean energy such as hydrogen separation by inorganic membrane, silica membrane reactors, modeling of membrane reactor performance, and separation processes. Satoru Nagasumi is Research Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Kyushu University in 2017. He started his research career to join the HTTR project at JAERI in 2017 for the research and development of the HTGR technology. He is mainly interested in nuclear physics of HTGRs. His current task is operation and maintenance of the HTTR, including evaluation of water overflow in the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. He has also joined the development

About the authors

xxi

of human resources to develop nuclear physics evaluation methods for HTGRs under the joint study between Kyushu University and JAEA. Shigeaki Nakagawa is Principal Researcher of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his BA degree from Kyushu University in 1986. He started his research career at JAERI in 1986 for HTGR safety. He then worked at the Los Alamos National Laboratory for 1 year in 1996 to develop the unattended spent fuel flow monitoring system for the HTTR. His research focuses on the areas of transient behavior coupling with neutronics and thermal hydraulics of HTGR during accident conditions to confirm its inherent safety. For this purpose, he plans the safety tests in the HTTR to demonstrate its inherent safety as HTGR. Tetsuo Nishihara is Deputy Director of International Cooperation and Social Environment Office, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Tohoku University in 1991. He started his research career at JAERI in 1991 for HTGR development. He carried out safety design for coupling the hydrogen production plan, and he conducted high-temperature long-term operation and safety demonstration test as the manager of the HTTR reactor engineering section in Department of HTTR. Hiroki Noguchi is Assistant Principal Researcher of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from the University of Saga in 2005. He started his research career at JAERI in 2005 for mechanical and chemical engineer. He then worked at Japan Science and Technology Agency from 2007 to 2009 and at German Aerospace Center from 2019 to 2020. His research focuses on the areas of iodine sulfur process for thermochemical hydrogen, which is one of the promising heat applications of hightemperature gas-cooled reactor. He has great interest in components development and demonstration of the iodine sulfur process. Naoki Nojiri is Manager of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his BA degree from Saitama University in 1994. He started his research career to join the HTTR project at JAEA in 1994 for the research and development of the HTGR technology. He is mainly interested in the quality assurance of nuclear installation. His current task is operational safety supervision of the HTTR. Hirofumi Ohashi is Deputy Director of Reactor Systems Design Department, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA.

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About the authors

He received his PhD degree from Yokohama National University in 1999. He started his research career at JAERI in 1999 for the research and development of the HTGR hydrogen production technology. His research focuses on the areas of safety design, system design, and safety analysis of HTGR. Masato Ono is Chief Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from the Kyoto Institute of Technology in 2010. He started his research career at the HTTR in 2010 for the research and development of the HTGR technology. After the Great East Japan Earthquake in 2011, he worked in the field of seismic evaluation of the HTTR. He has evaluated seismic integrity of the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. Akio Saikusa is General Manager of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Kyushu University in 1992. He started his research career to join the HTTR project at JAERI in 1992 for the research and development of the HTGR technology. He has been promoted to the operation and the project management of the HTTR. Nariaki Sakaba is Deputy Director of Strategy and Planning Office of Sector of Fast Reactor and Advanced Reactor Development, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD from Tohoku University in 2009. He started his research career at JAERI in 1991. He has worked for the HTTR project in the fields of design, construction, startup, and power-up testing, of the HTTR, and R&D on the IS process hydrogen production system. Before his current position, he held Deputy Director of Office of Strategy and International Affairs, Division Head of Hydrogen Application Research and Development Division, Group Leader of International Joint Research, and Principal Researcher of Policy Planning and Administration Department of JAEA. Koei Sasaki is Research Engineer of Department of Hydrogen and Heat Application Research and Development, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from the University of Fukui in 2016. He is in charge of maintenance of fuel handling machine in FBR, Monju from 2012 to 2017 at JAEA. Since 2018, he is responsible for research of the HTGR core materials. His main track record is an investigation of Fuel Clad Chemical Interaction on FBR cladding materials such as stainless steel, high chromium steel, and oxide dispersion strengthened steel. Chemical stability and irradiation stability of nuclear material is his main study field. His current research is to develop a coated fuel particle layer that has high retaining performance of fission product in HTGR.

About the authors

xxiii

Hiroyuki Sato is Group Leader of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Tokyo Institute of Technology in 2012. He started his research career at JAEA in 2006 for nuclear reactor system design and safety. His research focuses on the areas of safety and operability of advanced reactor systems. Especially, he is interested in the safety of nuclear system coupling to chemical plant. Hiroaki Sawahata is Chief Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his AS degree from National Institute of Technology, Ibaraki collage in 1994. He then joined HTTR project in 1994 and contributed to the achievement of first criticality, performance test, high-temperature continuous test, and safety demonstration of the HTTR. His current task is operation and maintenance of the HTTR, including an evaluation of protection against external event, management of beyond designbasis accident in the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. He is also interested in the development of operation and maintenance technology of HTGR. Taiju Shibata is Senior Principal Researcher of International Cooperation and Social Environment Office, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Ibaraki University in 2007. He started his research career at JAERI in 1993. He had stayed at Sevilla University in Spain from 2002 to 2003 to study about superplastic properties of ceramics. He is in charge of the promotion of international cooperation related to R&D and deployment of advanced reactors. His research focuses on the areas of graphite material of HTGR, specially graphite material properties, including neutron irradiation effects. He had worked as a member of Materials Project Management Board of VHTR system in GIF for a several years, and currently, he is a Japanese member of GIF Expert Group from 2018. He is engaged in the activities of TWG-GCR and TWG-SMR in the IAEA as a member of Japan. Yosuke Shimazaki is Chief Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Ibaraki University in 2008. He started his research career to join the HTTR project at JAEA in 2008 for the research and development of the HTGR technology. He is mainly interested in maintenance technologies and graphite components of HTGRs. His current tasks are nuclear material management of the HTTR and establishment of counter measure for beyond design-basis accident in the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements.

xxiv

About the authors

Atsushi Shimizu is Assistant Principal Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his BA degree from Shinshu University in 2001. He engaged in operation and maintenance of the HTTR. He contributed to the achievement of hightemperature continuous 50-day operation in 2010. In order to develop an evaluation method using PRA for HTGRs, he is investigating the failure rate database for HTGR-specific equipment using HTTR equipment operation and maintenance data. Since 2012, the safety design of the HTTR has been reviewed in order to meet the requirements of the new regulatory standards set by NRA, based on the lessons learned from the Fukushima Daiichi nuclear power plant accident. Masayuki Shinozaki is Director of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his BA degree from Kyushu Institute of Technology in 1988. He has been experienced in the design, construction, installation, inspection, testing, and in-service operation of the HTTR since 1988. He has been a general manager of operation and maintenance section of HTTR in 2010 and subsequently been a general manager of project management section in 2013. He successfully conducted the loss of cooling flow test in 2010. His current task is a representative of HTTR to correspond for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. Junya Sumita is Manager of Strategy and Planning Office, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Hokkaido University in 2008. He started his research career at JAERI in 1998 for fuel and material of high-temperature gas cooled reactor. He then worked at The University of Manchester in 2008. His research focuses on the graphite material for high-temperature gas-cooled reactor. He has been in his present position since 2018. Yukio Tachibana is Deputy Director General of HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from The University of Tokyo in 1987. He started his research career at JAERI in 1990. He has worked for the HTTR project in the fields of design, construction, startup, and testing, of the HTTR, design of HTGR commercial systems, and R&D on hydrogen production system by the IS process. He has interest in high-temperature strength of superalloys, high-temperature structural design guideline, and inelastic constitutive equations. Shoji Takada is Principal Researcher of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from the University of Hokkaido in 1999. He started his research

About the authors

xxv

career at JAERI in 1987 for the HENDEL project. He then worked in Juelich Research Center in Germany from 1999 to 2000. He joined the design study of gas turbine high-temperature reactor GTHTR300 and its R&Ds of high-performance helium gas compressor and magnetic bearing. He then joined the HTTR project in 2010 and successfully accomplished safety demonstration test of the HTTR. He has been promoted to a General Manager of reactor engineering section in 2015 and the General Manager of project management section in 2017. He is mainly interested in heat-transfer characteristics of HTGR and also in the development of operation and maintenance technology of HTGR. Kuniyoshi Takamatsu is Assistant Principal Researcher of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Kyushu University in 1999. He started his research career at JAERI in 2000 for the HTTR project. He also worked at the Argonne National Laboratory (ANL) from 2012 to 2013. He has experience with operation and maintenance of the HTTR, as well as R&D on thermal hydraulics, reactor dynamics, and radiation measurement. Tetsuaki Takeda has been a professor of Graduate Faculty of Interdisciplinary Research, Research Faculty of Engineering, Department of Mechanical Engineering, University of Yamanashi since 2008. He was graduated from Kobe University, majored in nuclear engineering in 1982, and received his doctoral degree at The University of Tokyo in 1997. He stayed at University of California Los Angeles (UCLA) as a visiting researcher from 1993 to 1994. He has transferred his field of academic goal from Japan Atomic Energy Agency to University of Yamanashi in 2008. His research field is the thermo-hydraulics related to not only nuclear energy but also renewable energy. During his service in JAEA, he performed the experiments and analyses of the safety studies regarding the VHTR systems and the nuclear hydrogen production systems developments. Afterward, he has also performed experiments and analyses regarding a ground source heat pump system, solar thermal collector system, heat utilization system using thermoelectric devices, and so on. He served Chairman of Power and Energy Systems Division of Japan Society of Mechanical Engineers (JSME). Hiroaki Takegami is Group Leader of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Nagoya University in 2003. He started his research career at Nagoya University in 2003 for metal joining process. He then worked at Nagoya Institute of Technology from 2005 for functional ceramic processing. His research focuses on the development of the IS process for the HTGR heat application. He is also interested in the strength evaluation method of ceramic components.

xxvi

About the authors

Nobuyuki Tanaka is Assistant Principal Researcher of Department of Hydrogen and Heat Application Research and Development, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from The University of Tokyo in 2014. He started his research career at JAERI in 2004, and he then worked at the Massachusetts Institute of Technology as Postdoctoral fellow in 2016. His research focuses on the research and development of thermochemical water-splitting hydrogen production method, the IS process. Especially, he has worked to improve process efficiency by using membrane separation technology. His research contributed to the success of the long-term continuous operation of hydrogen production of IS process. Daisuke Tochio is Assistant Principal Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from the University of Tsukuba. He started his research career at JAERI for the development of advanced nuclear reactor in 2003. His research focuses on the areas of thermal hydraulics in high-temperature gas-cooled reactor. His current task is operation and maintenance of the HTTR, including an evaluation of protection against BDBA, external event, and internal flooding in the HTTR facility for the conformity review by Nuclear Regulation Authority on the new regulatory requirements. He is also interested in two-phase heat transfer and fluid flow, especially vapor explosion. Takahiro Uesaka is Chief Engineer of Department of HTTR, HTGR Research and Development Center, Oarai Research and Development Institute, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his MS degree from Tokai University in 2005. He started his research career to join the Waste Treatment Facility in Nuclear Science Research Institute in JAERI in 2005. He developed lid type shielded container so that the modification of waste packages be removed to canned wastes easily from shielded containers, which meets the technical standards of disposal for intermediate-level wastes. He then received his MS degree from The University of Tokyo in 2013. He started his research career to join the HTTR project at JAEA in 2017. He is mainly interested in maintenance management of HTTR. His current task is operation and maintenance of specific facilities of the HTTR. Shohei Ueta is Assistant Principal Researcher of Reactor Systems Design Department, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD degree from Hokkaido University in 2009. He started his research career at JAERI in 1999 for the HTGR fuel. His research focuses on the areas of material science under the irradiation and manufacturing engineering for the extended burn up HTGR fuel for the practical HTGR. He is also interested in applied researches in advanced materials for upgrading performance and accident tolerance of the HTGR fuel. Since his research field

About the authors

xxvii

is closely related to radiation chemistry, he has great interest in characterization and modeling of the HTGR fuel and fission product behavior under irradiation. Xing L. Yan is Deputy Director of Reactor Systems Design Department, Sector of Fast Reactor and Advanced Reactor Research and Development in JAEA. He received his PhD from the Massachusetts Institute of Technology in 1990. Since then, he has been involved in various HTGR research and development works in the United States, Japan, and international programs. Currently, he manages the technical design of advanced HTGR plants and associated research, development, and innovation for fuel, material, components, safety technologies, reactor system concepts, and nuclear heat application systems in JAEA.

Preface of JSME Series in Thermal and Nuclear Power Generation

Electric power supply is a fundamental and principal infrastructure for modern society. Modern society uses power generation through heat. This series of books consists of eight volumes describing thermal and nuclear power generation, taking Japan as the example, and referring the other countries. The Volume 1 discusses how power supply is attained historically, focusing on the thermal and the nuclear power generation along with minimum-required scientific and technological fundamentals to understand this series of books. Then, the present status of the thermal and the nuclear power generation technique is displayed in detail in Volumes 2 through 8. The rehabilitation and reconstruction of Japan after World War II was initiated through the utilization of a large amount of coal for boilers of the thermal power plants. Meanwhile, environmental pollution caused by coal combustion became serious, and then oil was introduced to the boilers. Due to two worldwide oil crises and because of carbon dioxide issues, natural gas has also begun to be used for boilers. Current thermal power generation in Japan is based on coal and gas utilization. As a result of enough power supply, Japan has become one of the leading countries economically and technologically in the world. The thermal power technology that started from introducing technology from abroad has been transfigured Japan into one of the most advanced in the world through the research and development of Japanese industry, government, and academia during this process. Global warming related to excess carbon dioxide emissions has become a worldwide issue in recent years. Reducing carbon dioxide emission in thermal power generation is important to help cope with this issue. One direction is to change the fuel of a boiler from coal to gas that exhausts less carbon dioxide. Another important direction is to endeavor to enhance the thermal efficiency of coal thermal power plants as well as oil and gas. Many developing countries in the world need more thermal power plants in future. Although oil and/or gas thermal power plants may be introduced into these countries, it is supposed that coal thermal power plants will still be used due to economical reasons. Considering these situations, the publication of this series of books that displays and explains the developing history and the present status of the most advanced thermal power plants in Japan and other advanced countries is a timely planning for engineers and researchers in the advanced countries to pursue the further advancement and

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Preface of JSME Series in Thermal and Nuclear Power Generation

for engineers and researchers in developing countries to learn and acquire this knowledge. Nuclear power generation technology in Japan started after being introduced from abroad approximately 60 years ago. Then, it reached the matured nuclear power technology through untiring endeavors for research and development. However, nuclear power plants at the Fukushima Daiichi nuclear power station were heavily damaged by huge tsunamis caused by the Great East Japan Earthquake in 2011 to result in contamination in the large area around the power station. Taking measures to improve nuclear power reactors to be more robustly is currently underway by analyzing the factors that caused this serious situation. Technical vulnerability can be solved by technology. Nuclear power generation technology is one of the definite promising technologies that should be used in the future. The nuclear power generation is still expected as one of the main ways to supply electricity in the framework of the basic energy plan of Japan as well as thermal power generation. It implies that the construction of new power reactors will be required to replace the nuclear reactors that will reach their useful lifetime. Looking overseas, many developing countries are introducing nuclear power generation technology as a safe and economically excellent way to obtain electricity. Transfer of the nuclear power generation technology developed and matured in Japan to those countries is naturally the obligation of Japan. In these situations, the necessity of human resource development in the field of nuclear power generation technology in the developing countries, as well as in Japan, is beyond dispute. Thus, it is an urgent task to summarize the nuclear power generation technology acquired by Japan to provide it. The Power and Energy Systems Division (PESD) of the Japan Society of Mechanical Engineers is celebrating its 30th anniversary from establishment in 1990. This department is entrusted with handling power supply technology in mechanical engineering. Responding to the earlier-mentioned is truly requested. This task cannot be done by others but the PESD that is composed of leading engineers and researchers in this field in Japan. In view of these circumstances, summarizing Japan’s and other countries’ power generation technology and disseminating it not only in Japan but also overseas seems significantly important. So, it has been decided to execute this book series, publishing as one of the 30th anniversary events of the PESD. Authors of this book series are those who have engaged in the most advanced research and development for the thermal power and nuclear power generation in Japan and Canada. Their experience and knowledge is reflected in their writing. It is not an introduction of what others did, but living knowledge based on their own experiences and thoughts are described. We hope that this series of books becomes learning material that is not yet in existence in this field. We hope that readers acquire a way of thinking as well as whole and detailed knowledge by having this book series in hand.

Preface of JSME Series in Thermal and Nuclear Power Generation

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This series is the joint effort of many individuals, generously sharing and writing from their expertise. Their efforts are deeply appreciated. We are very thankful for the unbiased and heartful comments given from many reviewers to make this series superb. Special thanks should be given to Maria Convey and Sara Valentino of the editorial staff at Elsevier.

Editors in Chief Yasuo Koizumi Graduate School of Information and Engineering, The University of Electro-Communications, Chofu, Tokyo, Japan Mamoru Ozawa Kansai University, Takatsuki, Osaka, Japan March 16, 2020

Preface to Volume 5: High-Temperature Gas-Cooled Reactors

A high-temperature gas-cooled reactor (HTGR) is a nuclear reactor that can supply high-temperature heat energy of 750 C 950 C by using a spherical fuel coated with ceramics such as carbon and silicon carbide, inert helium gas as a coolant, and graphite as a moderator. HTGR can be used in various ways such as hydrogen production and process heat supply as well as power generation. At present, a light water reactor is mainly used for nuclear power generation, but HTGR is attracting attention mainly for the following two reasons: the first reason is that the nuclear heat should be used for fields other than power generation, because fossil fuels and renewable energy have a limit to meet the increasing energy demand in the future. The second reason is excellent safety. Among all reactor types, HTGR has excellent inherent safety and its safety has been demonstrated in experimental reactors. This volume provides the latest research on HTGR development and utilization, beginning with an analysis of the history of HTGRs. A detailed analysis of HTGR design features, including reactor core design, cooling tower design, pressure vessel design, I&C factors, and safety design, provides readers with a solid understanding of how to develop efficient and safe HTGR within a nuclear power plant. The following R&D on components for the design, safety review, and performance evaluation of the high-temperature engineering test reactor are described; fuel, core components and reactor internals, passive cooling system, intermediate heat exchanger, basic feature of air ingress during primary pipe rupture accident, and prevention technology for air ingress during primary pipe rupture accident. The targets of this volume are postgraduate course students, researchers, and engineers in the field of mechanical engineering, nuclear engineering, and chemical engineering. The areas are not only Japan but also the United States, Canada, China, and many other countries. We wish that readers acquire the way of thinking as well as whole and detailed knowledge from this volume. It is our great pleasure if they are able to enhance their understanding and talent. We also would like to offer our special thanks to the researchers and engineers who have been engaged in the research and development of HTGR for many years. Editors Tetsuaki Takeda and Yoshiyuki Inagaki2 1 University of Yamanashi, Kofu, Japan, 2Japan Atomic Energy Agency, Oarai, Japan April 2, 2020 1

Overview of high temperature gas-cooled reactor

1

Jin Iwatsuki, Kazuhiko Kunitomi, Hideaki Mineo, Tetsuo Nishihara, Nariaki Sakaba, Masayuki Shinozaki, Yukio Tachibana and Xing Yan Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

A high temperature gas-cooled reactor (HTGR) is a nuclear reactor that can supply high temperature heat energy of 750 C 950 C by using a spherical fuel coated with ceramics such as carbon and silicon carbide, inert helium gas as a coolant, and graphite as a moderator. At present, a light water reactor (LWR) is mainly used for nuclear power generation, but HTGR is attracting attention mainly for the following two reasons: the first reason is that the nuclear heat should be used for fields other than power generation, for example, the production of secondary energy such as synthetic fuel, methanol, and hydrogen, because fossil fuels and renewable energy have a limit to meet the increasing energy demand in the future and it is necessary to protect the global environment from the greenhouse effect and acid rain caused by carbon dioxide. The second reason is excellent safety. In the wake of the accidents at the Three Mile Island Unit-2 (TMI-2) in the United States and Chernobyl Unit-4 in the Soviet Union, research and development of a nuclear reactor that does not cause severe environmental pollution in any accident has been conducted including LWRs and fast reactors. Among all reactor types, HTGR has excellent inherent safety and its safety has been demonstrated in experimental reactors. The features of HTGR such as structure, heat application, and the history of development will be described.

1.1

Features of high temperature gas-cooled reactor

1.1.1 Structure and materials Fig. 1.1 shows the structure of HTGR constructed by Japan Atomic Energy Agency (JAEA), “the High Temperature Engineering Test Reactor (HTTR)” [1]. The structural and material differences of the fuel, coolant, moderator, etc. from the LWR used for power generation are described below. The details of the design, structure, and operation of the HTTR are described in Chapters 2 4.

High Temperature Gas-cooled Reactors. DOI: https://doi.org/10.1016/B978-0-12-821031-4.00001-4 © 2021 Elsevier Inc. All rights reserved.

2

High Temperature Gas-cooled Reactors

Helium coolant Stable at high temperature (no temperature limit) Ceramic fuel coa ng Retain radioacve material at 1600°C Graphite core structure Temperature limit: 2500°C

Coated fuel parcle

Fuel kernel High density PyC SiC

0.9 9 mm

Low density PyC

8mm 580mm 39m m

26mm

Fuel compact

Fuel assembly

Figure 1.1 Structure of HTTR [1].

1.1.1.1 Fuel LWR uses metal-clad fuels, while HTGR uses coated fuel particles (CFPs) with a diameter of about 1 mm, of which fuel kernel such as uranium oxide is coated in multiple layers with pyrolytic carbon and silicon carbide as shown in Fig. 1.1. These coatings contain the fission products (FPs) produced by the fission reaction of the fuel. The CRPs used in HTGR have excellent heat resistance, and it has been experimentally proven that the confinement function of the coating layer is not impaired even at a high temperature of 1600 C. Depending on the shape of the fuel, HTGR is classified into a prismatic type and a pebble bed type. In the prismatic type fuel, a fuel compact in which the fuel particles are sintered after mixing with graphite powder or a fuel rod in which the fuel compact is placed in a graphite sleeve is inserted into the graphite block. On the other hand, in the pebble bed type fuel, the CFPs are sintered into a spherical shape

Overview of high temperature gas-cooled reactor

3

with a diameter of about 6 cm after mixing with graphite powder. The HTTR is the prismatic type, but it is called a pin-in-block type because a fuel rod with a fuel compact in a graphite sleeve is inserted into a hexagonal columnar graphite block.

1.1.1.2 Coolant HTGR uses helium gas pressurized to 10 MPa or as the coolant. Helium gas is chemically inert and does not chemically react with fuels or structural materials even at high temperatures. In addition, there is no phase change in the operating temperature range, and helium gas has almost no nuclear effects such as neutron deceleration and absorption, and therefore does not affect the nuclear reaction of the core.

1.1.1.3 Moderator HTGR uses graphite as a moderator. Graphite is also a suitable core structural material because of its excellent properties such as low neutron absorption, high resistance to radiation, high heat resistance (sublimation temperature: about 3000 C), and high thermal conductivity. On the other hand, a large volume is required for the moderator because the slowing-down power is lower than that of water. As the graphite moderator also serves as a core structural material, the HTGR core size increases and the thermal power density decreases accordingly compared with LWR. As a result, the HTGR core has a large heat capacity, which contributes to excellent safety.

1.1.2 Heat application HTGR can supply very high temperature heat because it uses helium gas as the coolant, ceramics with high heat resistance as the fuel coating material, and has the core composed of graphite. The coolant outlet temperature of LWR is about 290 C 340 C, while that of HTGR is 750 C 950 C. Fig. 1.2 shows the ranges of temperature that can be supplied from HTGR and its potential heat applications. High temperature heat can be used for highly efficient gas turbine power generation. Large-scale hydrogen production by thermochemical water splitting process can be also carried out to make industrial uses such as steelmaking plant and transportation fuel. Process steam may be supplied in a wide range of temperature and pressure demanded by petroleum and chemical industries. In the 1960s, heat application of nuclear energy to the steelmaking process, namely nuclear steelmaking (NS), was proposed in response to the concern about the future shortage of coking coal and about dust pollution. The concept of NS is that reducing gases such as carbon monoxide and hydrogen are produced with heat of HTGR and supplied to a shaft furnace for steelmaking. The initial stage of this study was carried out by British Steel Corporation, Juelich Research Center, and EC. Research and development of HTGR in Japan was started for the purpose of NS.

4

High Temperature Gas-cooled Reactors

0

100

200

300

Temperature (°C) 400 500 600

700

800

900

1000

Thermochemical water splitting H2 production Gas turbine power generation Methane steam reforming

Styrene production

Oil shale recovery Petroleum refining

Seawater desalination, district heating LWR HTGR

Figure 1.2 HTGR heat supply temperature ranges and potential heat applications.

Due to the high reactor outlet temperature, HTGR power plant has higher thermal efficiency than those of LWR and the sodium fast reactor. While the thermal efficiency of LWR steam turbine power generation is less than 35%, the HTGR gas turbine power generation system can achieve high energy efficiency of around 46% at 850 C of the coolant. The high thermal efficiency brings low energy consumption and excellent energy security performance. Inherent safety features of HTGR also offer low construction cost by reducing safety equipment because HGTR can ignore the occurrence of severe accident and consequent core damage. The details of the heat application are described in Chapter 5, R&D on Commercial High Temperature Gas-Cooled Reactor.

1.1.3 Safety Helium gas does not react chemically with fuel and core structures so that hydrogen gas is not produced by chemical reaction of fuel element in accident like LWR as shown in Fig. 1.3 [1]. A large amount of water or air ingress can be eliminated by design of secondary water cooling system and reactor confinement building to prevent oxidation of fuel and core graphite material. HTGR does not need to consider the hydrogen explosion and vapor explosion. The ceramics CFP can bear very high temperature condition over 2200 C without any FP release as shown in Fig. 1.4 [1]. It can be reused under 1600 C by taking the safety margin. HTGR can be designed that the fuel temperature does not exceed 1600 C in any accident to prevent fuel damage. Therefore HTGR does not need to consider the core melt accident.

Overview of high temperature gas-cooled reactor

5

Figure 1.3 Safety feature of HTGR against hydrogen explosion [1].

Figure 1.4 Safety feature of HTGR against fission products release [1].

HTGR can remove the residual heat of the core indirectly because of optimized low reactor power density and graphite core structure. The core graphite has large heat capacity and high thermal conductivity. Since the forced cooling performance is gone in a loss of coolant accident, decay heat of fuel

6

High Temperature Gas-cooled Reactors

Figure 1.5 Safety feature of HTGR against accident management [1].

transfers to reactor vessel through the core graphite structure slowly by thermal conduction and emission as shown in Fig. 1.5 [1]. This performance restricts to rise the fuel temperature up to design limit of 1600  C. The HTGR does not need to consider the immediate accident management and to provide excess emergency safety system. A part of demonstration tests simulating this accident was performed with the HTTR as described in Section 4.5: Safety demonstration test.

1.1.4 Adaptability to environment The coated particle fuel of HTGR has high radiation resistance and the high FP containment performance. Average burnup ratio of the HTGR can be 120 GWd/t, which is three times higher than the metal cladding tube fuel like the LWR. The HTGR power generation system offers low amount of high-level radioactive waste in combination with the high burnup fuel and the high power generation efficiency. Preliminary estimation shows that the amount of the high-level radioactive waste can be reduced to one-fourth of LWR as shown in Fig. 1.6 [1]. The details of the fuel burnup and radioactive waste are described in Chapter 5, R&D on commercial high temperature gas-cooled reactor. HTGR also can supply high temperature heat and steam through secondary cooling system. This heat can be used as a heat source of chemical plant like hydrogen production plant as shown in Fig. 1.2. The replacement of HGTR from fossil fuel plant can contribute to reduce fossil fuel consumption and carbon dioxide emission. This contributes to resolve the global warming issue.

Overview of high temperature gas-cooled reactor

7

Figure 1.6 Low amount of high-level radioactive waste in HTGR [1].

1.2

History of research and development in world

Gas-cooled reactors were proposed as early as 1942, the year when the first nuclear pile (CP-1) went critical. In the mid-1940s, the design of a helium gas-cooled power reactor was proposed. In this design, the basic characteristics of advanced HTGR were established, namely, the use of helium gas as the coolant in a direct cycle gas turbine primary system, the selection of graphite as the moderator and core structural material, and the choice of uranium carbide and thorium carbide as the fissile and fertile materials, respectively, in a 235U/Th/233U fuel cycle. This concept was revised in the mid-1950s in the United Kingdom [2]. HTGR was expected to have the following advantages compared with other gascooled reactors represented by Magnox reactor. 1. 2. 3. 4.

HTGR can be downsized because of high thermal power density as a gas-cooled reactor. It has high thermal efficiency due to high outlet coolant temperature. From above, superior economics can be expected. In addition to power generation, it can be used as a heat source for various heat application processes.

At that time, the excellent inherent safety of HTGR was not emphasized as the reason for selecting a reactor, but rather superior economics was expected because no severe accident was experienced yet. In general, it is difficult to increase the cladding tube temperature above 650 C for a reactor with a fuel using a metal cladding tube such as LWR and a fast reactor. Therefore it is difficult to heat up their reactor outlet coolant temperatures above 600 C. In order to extract high temperature heat from the nuclear reactor,

8

High Temperature Gas-cooled Reactors

it was necessary to develop a cladding material, which has functions to contain fuel and prevent FP release and withstands a high temperature above 650 C. In 1956 CRPs were developed by the Harwell lab in the United Kingdom to solve this problem. With the development of CRPs, the Dragon project was proposed by the Harwell lab to the European Nuclear Agency in 1958. The Organization for Economic Co-operation and Development (OECD) accepted this proposal and established a program in which 12 European countries collaborated to develop HTGR. The 20 MWt experimental HTGR, Dragon, with a reactor outlet coolant temperature of 750 C was constructed in Winfrith, the United Kingdom. It was the first HTGR to attain criticality, but it did not generate electrical power. After the first criticality in 1964, it served as a most productive research tool for the development and soundness demonstration of fuel and graphite and accumulation of valuable experience in HTGR operation and maintenance, and the operation was terminated in 1976 [2]. After that, construction and operation of prismatic type and pebble bed type experimental reactors and power generation prototype reactors were carried out in the United States and Germany, respectively. Development of HTGR power generation plant started in early 1960s in the United States and Germany. They constructed and operated experimental reactor and demonstration reactor with steam turbine power generation system. But they terminated to construct a commercial reactor because of political and economic reasons. HTGR development has continued over half a century in world, during which the reactors listed in Table 1.1 have been built and operated in the world. The development at the moment is focused on the next-generation commercial designs, alternately called very high temperature reactor (VHTR), for a range of economic applications, of which one that promises to exploit the highest potential of the HTGR is the production of hydrogen [3]. In the United States, General Atomics (GA) constructed a power generation prototype reactor (thermal output: 115 MWt, coolant outlet temperature: 725 C), Peach Bottom, from the late 1950s. The fuel kernels, originally coated with one layer of pyrolytic carbon (PyC), were prismatic and used uranium and thorium carbides as fuel. It was the first HTGR to generate electrical power (electrical output: 40 MWe) with a steam turbine, and operated successfully from 1967 to 1974. It was also operated as a test bed for advanced coated particle graphite matrix fuels for large HTGRs. Following the Peach Bottom, GA built a power generation prototype reactor, Fort St. Vrain (FSV), of which thermal and electrical outputs were 842 MWt and 330 MWe, respectively, and a reactor outlet temperature of coolant was 775 C. Major features of the FSV are use of a prestressed concrete reactor vessel (PCRV), once-through modular steam generators with integral superheaters and reheaters, steam-driven axial flow helium circulators, and hexagon-shaped graphite fuel elements incorporating improved, carbon and carbide CFPs [4]. The FSV attained criticality in 1974. After experiencing initial failures such as water ingress into the primary coolant system and fluctuations in the temperature at the core outlet, it started commercial operation with 70% output from 1979 and achieved 100%

Table 1.1 HTGRs constructed in world [3]. Test HTGRs

Country Period of operation Reactor type Thermal power, MWt He coolant outlet temp.,  C Coolant pressure, MPa Electrical output, MW Process heat output, MW Process heat temp.,  C Core power density, W/cm3 Fuel particle Kernel coating

Prototype HTGRs

Dragon

AVR

HTTR

HTR-10

Peach Bottom

FSV

THTR-300

UK (OECD) 1963 76 Tube 21.5 750 2

Germany 1967 88 Pebble 46 950 1.1 13

Japan 1998 present Prismatic 30 950 4.0

China 2000 present Pebble 10 700 3.0 2.5

USA 1967 74 Tube 115 725 2.25 40

USA 1976 89 Prismatic 842 775 4.8 330

Germany 1986 89 Pebble 750 750 3.9 300

2 UO2 TRISO

8.3 ThC2 BISO

6.3 (Th/U, Th)C2 TRISO

6.0 (Th/U)O2 BISO

14 UO2 TRISO

2.6 (Th/U,U)O2, C2 BISO & TRISO

10 863 2.5 UO2 TRISO

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High Temperature Gas-cooled Reactors

output in 1981 with a net overall efficiency of approximately 40%. The operation of the FSV terminated in 1989 because the operation rate was not always good. In Germany, 15 municipal electricity companies formed a subsidiary company to operate Arbeitsgemeinschaft Versuchsreaktor (AVR), which was constructed as an experimental power station for pebble-bed reactors with an additional purpose of testing fuels. The thermal and electrical outputs with a steam turbine were 46 MWt and 13 MWe, respectively, and a core outlet temperature of coolant was 950 C. The core was fueled with around 100,000 graphite pebbles containing CFPs. During operation, the 6-cm-diameter pebbles were continuously withdrawn from the bottom of the reactor core and other pebbles were added at the top of the core so that circulation takes place. After several passes through the core, the spent pebbles, which were removed from the cycle when their target burnup was achieved, were replaced by new pebbles. The coolant flowed upward through the pebble bed and then across the steam generator tubes, which were located above the core in the steel reactor vessel [5]. The AVR attained criticality in 1966, started generating electricity in 1967, reached full power in 1968, and terminated operation in 1988. The AVR-U project, which modifies the AVR and supplies helium gas at 950 C to the steam reformer and intermediate heat exchanger (IHX) located outside the reactor for nuclear heat application experiments, was studied; however, it was canceled due to budgetary reasons. A pebble bed type and power generation prototype reactor, Thorium Hochtemperatur Reaktor (THTR-300), was sponsored by the Federal Republic of Germany (FRG) and Nordrhein Westfalen. The thermal and electrical outputs with a steam turbine were 750 MWt and 300 MWe, respectively, and a reactor outlet temperature of coolant was 750 C. The reactor vessel was a PCRV, and the fuel particles used were uranium and thorium oxide. Its construction began in 1971 but, primarily because of changing licensing requirements, was not completed until 1984. The reactor plant was connected to the electrical grid of the HochtemperaturKernkraftwerk GmbH (HKG) utility in 1985. In 1989 the decision was made to permanently shut down the THTR-300 for sociopolitical reasons, not because of technical difficulties associated with the plant. These sociopolitical reasons were enacted by an application by HKG for early decommissioning based on a projected shortfall in funding and contractual changes in the allocation of decommissioning costs between the FRG, Nordrhein Westfalen, and HKG that would take effect upon the termination of the demonstration phase in 1991 [6]. Asia is home to the newest builds, the HTTR, and high temperature reactor 10 MWt (HTR-10), which are prismatic and pebble-bed designs, respectively, and are operational today. The HTTR rated at 30 MWt has a reactor outlet coolant temperature of 950 C and allows for 863 C process heat output. Such high temperature capability is compatible with modern process technologies and widens market roles of the reactor, as shown in the recent commercial designs. In China, a demonstration reactor plant (HTR-PM) is under construction (connection to grid is planned in 2020), and commercial reactor plant programs are practically in progress. Sponsored mainly by the U.S. Department of Energy (DOE), a research and industrial group proposed the gas turbine modular helium reactor (GT-MHR) in

Overview of high temperature gas-cooled reactor

11

1994 [7]. The design is based on a 600 MWt and 850 C prismatic core reactor that is passively safe and employs gas turbine power conversion at thermal efficiency approaching to 50%. The cost of electricity generation was shown competitive to other generation options [8]. GA has since continued the GT-MHR development in cooperation with partners in the Russian Federation. The Japan Atomic Energy Agency (JAEA) has offered the GTHTR300C, a gas turbine high temperature reactor of 300 MWe for cogeneration of variable rated of electricity and hydrogen [9]. The GTHTR300C used a 600 MWt prismatic core reactor with outlet coolant temperature of 950 C to power a direct cycle gas turbine for electricity generation and thermochemical process for hydrogen production. South Africa has been developing the 400 MWt pebble-bed modular reactor (PBMR) with reactor outlet coolant temperature of 900 C for the production of electricity and hydrogen and for other process-heat applications [10]. In 2001 Generation IV International Forum (GIF) of 10 member countries endorsed 6 nuclear systems that can be licensed, constructed, and operated by the year 2030 and which will deliver affordable energy products while satisfactorily addressing the issues of nuclear safety, waste, and proliferation [11]. Recognizing the VHTR to be deployable in the near future and exceptionally suitable not only for electricity generation but also for hydrogen production and industrial heat applications, the U.S. DOE has placed the Generation-IV priority on the VHTR. This led the United States to the creation of the next-generation nuclear plant (NGNP) program to demonstrate commercial high-efficiency generation of electricity and hydrogen. In the world now, policy, economy, society, and environment are becoming unstable; energy security and global warming restriction are the common subjects to be solved. “Nuclear energy,” in particular, is reviewed in light of its sustainability, cleanliness, and diversity, although some countries became “Away from nuclear,” many countries are promoting nuclear energy renewed development and/ or its introduction, through enhancement of safety under severe accident conditions. Introduction of HTGR is planned in Poland. Japan supports the plan with international cooperation [12]. The United Kingdom, Canada, and the United States are planning to introduce SMR including HTGR.

1.3

History of research and development in Japan

In Japan, Committee on Utilization of Nuclear Energy, the Iron and Steel Institute of Japan, was established in 1968. The VHTR round-table conference of the Atomic Energy Commission described in the report compiled in 1971 that it was important to advance immediately the full scale research and development of NS from the standpoints of the stable reservation of the energy security by converting the dependence to coking coal into nuclear energy, that is diversification of energy supply, and promotion of the technical innovation by development of independent technology. Based on the investigations by the committees, Research and Development of Nuclear Steelmaking was inaugurated in 1973 as a national R&D

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High Temperature Gas-cooled Reactors

project by the Agency of Industrial Science and Technology, Ministry of International Trade and Industry. The Engineering Research Association of Nuclear Steelmaking (ERANS), an industrial consortium of companies and institute, promoted this R&D program from 1973 to 1980 [13]. Since the 1970s, there was no shortage of coking coal, so the R&D of NS did not move to the next stage. The project by the ERANS was carried out under the following subthemes: (1) total system design, (2) high temperature heat exchanger, (3) heat-resistant superalloys, (4) high temperature thermal insulator, (5) reducing gas production unit, and (6) reduced iron production unit. They were the principal constituents in establishing the fundamental engineering technologies for designing, constructing, and operating a pilot plant, which was to incorporate the 50 MWt experiment VHTR to be developed by Japan Atomic Energy Research Institute (JAERI, reorganized in October 2005, at present JAEA) [14]. As for the total system, the conceptual designs of the pilot plant with the 50 MWt VHTR and the commercial prototype plant with the 500 MWt VHTR were performed. Fig. 1.7 shows a conceptual process flow scheme of a pilot plant with 50 MWt VHTR, FM-50 system. The secondary helium gas transports thermal energy from the IHX to the process plant, consisting of five main components, that is, a steam reformer (RF), a steam heater (SH), a reducing gas heater (RGH), and two steam generators (SG). In the reducing gas production unit, the vacuum residue is decomposed into light hydrocarbons and pitch by the high temperature steam cracking process. The light hydrocarbon, which includes naphtha and cracked gas, is desulfurized and fed to the steam reformer, where the light hydrocarbon absorbs heat from the secondary helium in the presence of a catalyst and gets changed into the reducing gas. On the other hand, pitch is converted to reducing gas by means of water-gas reaction. The reducing gases produced by these two processes are mixed and purified in the acid gas removal section. After being purified, the gas is heated up to 850 C in the reducing gas heater and fed to the high pressure shaft furnace [13]. Fig. 1.8 shows the conceptual flow diagram of the prototype plant. Two heat utilization systems, RF and SH systems, which have 253 MWt IHXs, are connected to the 500 MWt VHTR. The RF system supplies the reducing gas produced by steam reforming of the methane manufactured from coal liquid to the shaft

Figure 1.7 FM-50 system process scheme [13].

Overview of high temperature gas-cooled reactor

13

Figure 1.8 Process scheme of 500 MWt prototype nuclear process heat plant [13].

furnace. Meanwhile, the SH system decomposes vacuum residue with high temperature steam and performs production of methane, naphtha, etc. and power generation [13]. As for the high temperature heat exchanger, basic examinations such as improvement in heat transfer were carried out and a design code of high temperature heat exchanger serviced in 1000 C was created. Then, the base technology for heat exchange was established by manufacture of a 1.5 MWt IHX and a high temperature helium test loop and around 2000 h operation in 1000 C. This result was reflected in the design and manufacture of the 10 MWt IHX of the HTTR. The new heat-resistant superalloys used for the heat exchanger tube of IHX in helium of 1000 C were developed [15], and R&D were continued by JAERI after the end of this project. In the meantime, the main specifications of HTGR were revised by JAERI toward early construction, and the coolant temperature was changed into 950 C from 1000 C. When construction of the HTTR began in 1991, the superalloy applicable to the HTTR was only Hastelloy XR developed in JAERI because the material data of the developed new superalloys were insufficient for design and fabrication. As for the high temperature thermal insulator, the performance of aluminasilica fiber insulator such as resilience was improved and the demonstration test was carried out with the high temperature helium test loop. This insulator was applied to the IHX, hot gas ducts, etc. of the HTTR. As for the reducing gas production unit, a heat exchange developed and the demonstration operation was performed for about 3000 h. The reducing gas production technology from vacuum residue was established. The cold partial model tests of the shaft furnace and research of reducing condition were carried out in parallel for the reduced iron production unit, and then the shaft furnace plant was designed. This NS project clarified the future vision of nuclear heat application. As for installation of chemical plants such as reducing gas production plant, etc. closed to the VHTR, however, there were some problems such as safety design against fire and explosion of combustible gas and disclosure of poisonous gas, a measure against tritium permeation from VHTR to the chemical plant and so on. These

14

High Temperature Gas-cooled Reactors

problems are continually researched by JAEA. The safety design and integration technology to connect chemical plants to HTGR (and/or VHTR), and new NS system design is described in Chapter 5, R&D on Commercial HTGR. The development of HTGR aimed at the process utilization of heat generated by the reactor as a heat source not only for power generation but also for ironmaking, and chemical industry was taken up as the theme of JAERI. In this project, the design of the experimental reactor and the development of related component technology have been carried out. As for development of related component technology, R&D in all fields required for reactor design and construction, including R&D on reactor physics, coated particle fuel, heat-resistant materials, high temperature equipment, have been advanced using the Helium Engineering Demonstration Loop (HENDEL), the Very High Temperature Reactor Critical Assembly (VHTRC), the Oarai Gas Loop-1 (OGL-1), etc. in JAEA. Since 1985, JAEA started the HTTR project, which is the first HTGR in Japan. JAEA got a construction permission of the HTTR in 1990 and started construction in 1991. The HTTR achieved the first criticality in 1998 and the world’s first operation of 950 C reactor outlet coolant temperature in 2004. JAEA has thus completed the world-highest-level HTGR technologies. Furthermore, JAEA conducted long-term high temperature operation (950 C/50 days operation) to demonstrate the capability of high temperature heat supply and the loss of forced cooling (LOFC) test at reactor power of 30% to demonstrate the inherent safety feature of HTGR in 2010. The LOFC test simulates the severe accident in which

Figure 1.9 History of HTGR development in Japan [1]

Overview of high temperature gas-cooled reactor

15

the reactor coolant flow is reduced to zero and the reactor scram is block. The test result shows that reactor shuts down by Doppler Effect and keeps stable condition without any operation management. JAEA has accumulated useful data for development of future commercial HTGR system through design, construction, and operation of the HTTR. History of HTGR development in Japan is shown in Fig. 1.9. Toward restart of the HTTR after the Great East Japan Earthquake that occurred on 11 March 2011, the official safety review meeting of the application of the HTTR to the Nuclear Regulation Authority (NRA) to confirm the adjustability to the new regulation standard has been carried out since 2014, by the NRA. After restart of the HTTR, it is planned to implement the LOFC test that simulates the loss of all AC power. “Chapter 3. Promotion of technology development, 2. Technical challenges to be addressed” of the 5th Energy Basic Plan, which was decided by the Cabinet in July 2018, mentions that “Under international cooperation while looking at overseas market trends, GOJ also facilitates R&D of nuclear technologies that serves the safety improvement of nuclear use, such as HTGRs which are expected to be utilized in various industries including hydrogen production and which has an inherent safety.” The industry, academic body, and government council established by the Ministry of Education, Culture, Sports, Science and Technology (MEXT) are developing strategies for commercializing HTGRs and deploying Japanese HTGR technologies overseas. A cooperation policy and domestic organization structure are consideration for an experimental and commercial reactor in Poland.

References [1] T. Nishihara, et al., Excellent Features of Japanese HTGR Technologies, Japan Atomic Energy Agency, JAEA-Technology 2018-004, 2018. [2] M.T. Simnad, The early history of high-temperature helium gas-cooled nuclear power reactors, Energy 16 (1991) 25 32. [3] X. Yan, et al., Nuclear Hydrogen Production Handbook, CRC Press, 2011. [4] A.L. Habushu, A.M. Harris, 330-MW(e) Fort St. Vrain high-temperature gas-cooled reactor, Nucl. Eng. Des. 7 (1968) 312 321. [5] K. Kru¨ger, et al., Preparation, conduct, and experimental results of the AVR loss-ofcoolant accident simulation test, Nucl. Sci. Eng. 107 (1991) 99 113. [6] J.M. Beck, L.F. Pincock, High Temperature Gas-Cooled Reactors Lessons Learned Applicable to the Next Generation Nuclear Plant, Idaho National Laboratory, INL/EXT10-19329 Revision 1, 2011. [7] U.S. DOE, Evaluation of the Gas Turbine Modular Helium Reactor, U.S. DOE-GTMHR-100002, 1994. [8] P.M. Williams, et al., MHTGR development in the United States, Prog. Nucl. Energy 28 (1994) 265 346. [9] X. Yan, et al., GTHTR300 design variants for production of electricity, hydrogen or both, in: Proceedings of the OECD/NEA 3rd Information Exchange Meeting Nuclear Production Hydrogen, OECD/NEA, 2005, pp. 121 139.

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High Temperature Gas-cooled Reactors

[10] D. Matzner, PBMR project status and the way ahead, in: Proceedings of the 2nd International Topical Meeting on High Temperature Reactor Technology, Beijing, China, 2004. [11] U.S. DOE Nuclear Energy Research Advisory Committee and the Generation IV International Forum, A Technology Road Map for Generation IV Nuclear Energy Systems, GIF-002-00, 2002. [12] Japan-Poland Foreign Ministers’ Meeting, The two ministers signed the “Action Plan for the Implementation of the Strategic Partnership between the Government of Japan and the Government of the Republic of Poland”, where encouraging cooperation toward R&D of HTGR between the JAEA and NCBJ, ,https://www.mofa.go.jp/press/release/ press4e_001593.html., May 18, 2017. [13] K. Tsuruoka, et al., Design study of nuclear steelmaking system, Trans. Iron Steel Inst. Jpn. 23 (1983) 1091 1101. [14] K. Shimokawa, Present status of research and development of nuclear steelmaking in Japan, Trans. Iron Steel Inst. Jpn. 19 (1979) 291 300. [15] R. Tanaka, et al., Development of superalloys for intermediate heat exchanger tubes in national research and development program of nuclear steelmaking, Tetsu-to-Hagane 68 (1982) 226 235.

Design of High Temperature Engineering Test Reactor (HTTR)

2

Yusuke Fujiwara, Minoru Goto, Kazuhiko Iigaki, Tatsuo Iyoku, Hai Quan Ho, Taiki Kawamoto, Makoto Kondo, Kazuhiko Kunitomi, Keisuke Morita, Satoru Nagasumi, Shigeaki Nakagawa, Tetsuo Nishihara, Naoki Nojiri, Masato Ono, Akio Saikusa, Nariaki Sakaba, Taiju Shibata, Yosuke Shimazaki, Atsushi Shimizu, Masayuki Shinozaki, Junya Sumita, Yukio Tachibana, Shoji Takada, Daisuke Tochio, Takahiro Uesaka and Shohei Ueta Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

The HTTR is the first and the only HTGR in Japan, which is installed at the Oarai Research and Development Institute of JAEA. In this chapter, the design details of the HTTR are described as follows: 1. Overview of HTTR design features The general philosophy and specific aspects of the HTTR design and its major characteristics together with history, back ground and objective. 2. Nuclear design Nuclear characteristics, such as reactivity control, power distribution, reactivity coefficients, etc., were evaluated. The reactor can be shut down safely by either the control rods or the reserved shutdown system. The reactor possesses power suppressing safe characteristic due to negative reactivity coefficients. 3. Core thermal-hydraulics The maximum fuel temperature at normal operation and anticipated operation occurrences and the coolant flow rate distribution were evaluated. The safety analysis clarified the maximum fuel temperatures at normal operation and at any anticipated operation occurrences do not exceed the each design limit. 4. Graphite components The design criteria for the graphite components have been developed by JAEA because there were no available design criteria for brittle materials such as graphite. An outline of quality control specified in the design criteria is also described. 5. Metallic components Hastelloy XR was developed for the very high temperature structures such as the heat transfer tubes of the intermediate heat exchanger (IHX) through optimizing or lowering contents of several elements of Hastelloy X. The high temperature structural design guidelines for Hastelloy XR, 21/4Cr1Mo steel, SUS316 and SUS321 austenitic stainless steels and 1Cr0.5MoV steel were established based on the existing design guidelines. 6. Core components and reactor internals Design requirements were settled and extensive design efforts and R&D were performed in order to maintain the integrity of the core components and reactor internals. Assessment of fuel integrity was made comprehensively and the fuel integrity was confirmed as to normal operating and transient conditions of the HTTR. The structural integrity of the graphite components was also ascertained by performing the analytical evaluations. High Temperature Gas-cooled Reactors. DOI: https://doi.org/10.1016/B978-0-12-821031-4.00002-6 © 2021 Elsevier Inc. All rights reserved.

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High Temperature Gas-cooled Reactors

7. Seismic design The validity of the core seismic analysis code called SONATINA-2V was confirmed. Impact load characteristics for the dowelsocket and keykeyway systems and core support posts were clarified quantitatively. The structural integrity of the graphite components against earthquake was also confirmed throughout the seismic tests and the analysis. 8. Cooling system The schematic flow diagram of the primary and secondary cooling systems and design requirements of key components such as the primary pressurized water cooler, the IHX, the helium gas circulator and the concentric hot gas duct are described. 9. Reactivity control system A design requirement of control rods used in the very high temperature environment is described. Alloy 800H was chosen for the metallic parts of the control rods, because the maximum temperature of the control rods reaches about 900  C at reactor scrams. The design guideline for the HTTR control rod was made based on ASME Code Case N-47-21. Observing the guideline, temperature and stress analysis were conducted; it can be confirmed that the target life of the control rods of 5 years can be achieved. 10. Instrumentation and control system The design outline of the instrumentation, control equipment and safety protection systems and the schematic diagram of control system were described. Their performance such as the in-core temperature monitoring, the reactor power control, the reactor outlet coolant temperature control, etc. was clarified at the rise-to-power test. 11. Containment structures The design specifications of the reactor containment vessel, the service area and the emergency air purification system are described. The measured leakage rates of the reactor containment vessel were enough less than the specified leakage limit of 0.1 %/d confirmed during the commissioning tests and annual inspections. The service area was kept in a way that the design pressure becomes well below its allowable limitation by the emergency air purification system. 12. Other systems The design specifications of the auxiliary helium systems such as the helium purification, the helium sampling, the helium storage and supply systems, and the fuel handling and storage system. 13. Safety design A safety design philosophy of the HTTR considering the major design features of HTGR is described. The strategy of defense in depth was implemented so that the safety engineering functions such as control of reactivity, removal of residual heat and confinement of fission products shall be well performed to ensure safety. The analytical codes for the safety evaluation are also introduced.

2.1

Overview of HTTR design features

2.1.1 Introduction JAEA has carried out the research and development on HTGR and hightemperature heat applications since the 1960s. Based on the Long-term Program for Research, Development and Utilization of Nuclear Energy revised by the Atomic Energy Commission of Japan in 1987, the construction of the HTTR was determined and initiated at the Oarai Research and Development Institute of JAEA in 1991. The HTTR attained the first criticality on November 10, 1998, and achieved the full power of 30 MW with the reactor outlet coolant temperature of 850 C on

Design of High Temperature Engineering Test Reactor (HTTR)

19

December 7, 2001. After series of safety demonstration tests, it will be used as heat source of a hydrogen production system [1]. The purpose of the HTTR project is to establish HTGR and nuclear heat utilization technologies and to carry out basic researches on high temperature irradiation. Recently, JAEA has stressed the importance of research and development on the hydrogen production considering significance of hydrogen as an energy carrier for energy security and prevention of global climate change. In order to achieve early deployment of a hydrogen society, which is beneficial for human health and the environment, JAEA is acting as the prominent organization for the HTGR and hydrogen production technologies, and the HTTR is used as a keystone for the development of these technologies. The HTGR technology is now being established through the accumulation of the operation experience and various test data of the HTTR. As for the nuclear heat application research, development of a hydrogen production system to be connected to the HTTR has been conducted, aiming at realizing the hydrogen production using the nuclear heat. Especially, hydrogen production by the water splitting iodinesulfur (IS) method that has been developed for a long time by JAEA is seen as the efficient and ultimate clean system, and thus the IS process is considered as the primary candidate to be connected to the HTTR. The GTHTR300 system for generating electricity with high thermal efficiency of about 46%, combining a HTGR and a gas turbine generator system, is also being developed. The basic researches include the development of new materials and the development of high temperature in-core instrumentation technologies. The HTTR project enables the HTGR and related nuclear heat application technologies to be deployed. As a result, the nuclear heat will be applicable to various fields such as chemical industries, which currently emit large amount of carbon dioxide. The hydrogen production using nuclear heat will completely cut the emission of carbon dioxide and make it possible to realize an ultimate clean hydrogen society. The system for generating electricity with high thermal efficiency of about 46% is expected to be a safe and economically competitive system. The HTGR and related heat application technologies can contribute to the environmental preservation by reducing dependence on fossil fuels and effectively using the nuclear energy.

2.1.2 History and future plan of HTTR project Fig. 2.1 shows the history of the HTTR project. For establishing and upgrading HTGR technologies, JAEA decided to construct the high temperature engineering test reactor based on the Long-term Program for Development and Utilization of Nuclear Energy, which was revised in 1987. JAEA obtained the installation permit for the HTTR from the government in November 1990. JAEA received the first approval of the HTTR design and construction methods from the Science and Technology Agency (STA) in January 1991 and started the construction work for the HTTR in March 1991. Excavation was completed in August 1991 and the construction of the concrete base mat was completed in May 1992.

20

High Temperature Gas-cooled Reactors

Figure 2.1 History of HTTR project.

The reactor containment vessel was installed and passed proof pressure and leakage tests successfully in November 1992. The reactor pressure vessel (RPV), intermediate heat exchanger (IHX), primary helium circulators, and primary-pressurized water cooler (PPWC) were installed in the reactor containment vessel in 1994. The first pressure test for the primary cooling system was carried out successfully in October 1995. Construction of the reactor building was completed by closing the temporary opening for carrying in large components in December 1995. JAEA obtained the uranium material for the first loading fuel in 1994, and manufacture of the fuel rods was completed in October 1997. Assembly of fuel elements was carried out in the reactor building and was completed in December 1997. Comprehensive and functional tests for each system were started in October 1996, and several malfunctions such as a temperature rise of the upper concrete biological shield of the RPV were found. The improvements were finished by March 1998 [2]. Preparation for the first fuel loading of the HTTR was started in April 1998. During this preparation, some design problems were found in the fuel handling machine. After countermeasures against these problems were finished, the first fuel loading was started on July 1, 1998. The first criticality was achieved in a core of 19 columns at 14:18 on November 10, 1998. Loading of all 150-fuel assemblies was completed on December 16. The criticality tests of the full-loaded core were carried out in January 1999. The criticality tests confirmed the core characteristics of the HTTR. Improvement of the HTTR reactor system and confirmation tests was carried out until August, then the reactor was considered to be ready for the rise-to-power test. The rise-to-power test was started on September 28, 1999, and the HTTR achieved the full power of 30MWat a reactor outlet coolant temperature of about 850 C on December 7, 2001. The operation permit of the HTTR was issued on March 6, 2002 from the Ministry of Education, Culture, Sports, Science and Technology (MEXT). The following extensive tests using the HTTR were started in

Design of High Temperature Engineering Test Reactor (HTTR)

21

FY 2002. The high temperature test operation was carried out and the reactor outlet coolant temperature of 950 C was achieved on April 19, 2004. The long-term high temperature operation was successfully completed by using the HTTR at the rated power 30 MW by maintaining the reactor outlet temperature at 950 C for 50 days from January to March, 2011, subsequently after the longterm rated power operation at 850 C for 30 days. The potential of stable hightemperature heat supply to the heat utilization system such as the planned HTTR-IS hydrogen production system was demonstrated. The technical basis of HTGRs was established through the evaluation of the test data [3] concerning the core physics [4], the fission product (FP) retention performance of coated fuel particles (CFPs) [5], the impurity control in coolant helium [6], and the operation and maintenance technology [7]. The safety demonstration tests have been carried out to verify the inherent safety features of the HTGR. It was revealed that the negative feedback effect of the core brings the reactor power lower as the core temperature rises, finally to a safe and stable power level without a reactor scram through the tests simulating the anticipated operation occurrences (AOOs). The control rod withdrawal tests and the coolant flow reduction tests have been successfully carried out to verify the safety evaluation method to reveal the inherent safety of HTGRs [8]. As a result, a loss of forced cooling (LOFC) test using the HTTR was licensed by regulatory authority in 2006 to verify the inherent safety of HTGR under the condition of LOFC while the reactor shutdown system is disabled. The LOFC test simulates the anticipated transients without scram (ATWS), including a LOFC in the core with deactivation of all trains of cooling in vessel cooling system (VCS). The first LOFC test was carried out at thermal power of 30% by using the HTTR in December 2010 [9]. Subsequently, the test of the LOFC with a partial loss of cooling in the VCS was carried out at thermal power of 30% in January 2011. A part of the tests is carried out as an international joint research for safety of innovative reactors organized by the OECD/NEA, the OECD/NEA LOFC project, in March 2011. JAEA has carried out the inspections of the structures and systems, etc. in parallel with the seismic analysis by using the actual seismic waves measured on the reactor during the earthquake on March 11, 2011, which are needed before restarting the reactor [10]. A new regulation standard issued on December 18, 2013, by Nuclear Regulation Authority (NRA). JAEA submitted the evaluation results satisfying the new regulation standard to the NRA on November 26, 2014, which has been reviewed by the NRA. Fig. 2.2 shows the R&D of HTGR using HTTR and Introduction Plan.

2.1.2.1 Evaluation of reactor performance Based on the HTTR operational data, the HTGR reactor performance has been evaluated and analytical computer codes are verified or modified for predicting realistic reactor performance under steady-state and operational transient conditions. The evaluation is focused on: (1) core physics in relation with thermal response and control system, (2) thermal analysis for fuel, reactor internals, and high temperature

22

High Temperature Gas-cooled Reactors

Figure 2.2 Present and future plan of the HTTR project.

components, (3) fuel performance on FP release and degradation of the coating layers to contain the FPs, (4) structural integrity of reactor internals and high temperature components, (5) decay heat and residual heat removal characteristics. The fruits from the HTTR operational data and their evaluation are expected to be utilized for the design of the future Japanese advanced HTGR.

2.1.2.2 Safety demonstration test It is well known that safety of medium-size HTGRs can be maintained even in the case of no forced cooling systems functioned with failure of reactor shutdown because of its strong negative feedback, large heat capacity of the core, etc. It is of great importance and one of the best ways for the wide public acceptance to demonstrate such inherent safety of the HTGR using an actual HTGR. It is, therefore, planned in the HTTR to conduct a safety demonstration test. The safety demonstration test may not be enough to convince the public to accept the safety of the HTGR; however, it is helpful from the standpoint of the public acceptance. The obtained test results are useful for validation of HTGR safety analysis codes. The safety demonstration test is divided into two phases. The first phase test, which simulates anticipated operational occurrences, includes primary coolant flow reduction test and a control rod withdrawal test at power operation. In the primary coolant flow reduction test, coolant flow rate is reduced by running down one and two gas circulators out of three. This test demonstrates that rapid decrease of coolant flow rate brings reactor power to stable level by the negative reactivity feedback of the core without a reactor shutdown. Due to these tests, the basic characteristics of the HTTR are confirmed and analytical codes for reactor transients are validated. The second phase test simulating accident

Design of High Temperature Engineering Test Reactor (HTTR)

23

conditions will be conducted after completion of the first phase test. The second phase test contains LOFC test and VCS stop test.

2.1.2.3 Development of process heat application system To enhance the nuclear energy application for heat process industries, JAEA has continued extensive efforts for development of hydrogen production systems using nuclear heat from an HTGR. In the hydrogen production by well-established steam reforming of natural gas, the emission of CO2 is inevitable because natural gas of methane is used as feed gas. Therefore the final goal of the hydrogen production system using HTGR is to produce hydrogen from water without emission of CO2. For this purpose, the thermochemical iodinesulfur process is under study. After the successful completion of laboratory-scaled test, an engineering basic test is under way aiming to develop the process control technique. Continuous hydrogen production tests with the test facility made of industrial materials such as metal and ceramics are currently under way. A process heat application system will be coupled to the HTTR in the future, and hydrogen will be produced directly by using nuclear energy.

2.1.3 Major design features of HTTR The major specifications of the HTTR are summarized in Table 2.1. The HTTR consists of the reactor core, a reactor cooling system, engineered safety systems, instrumentation, and control system. Figs. 2.32.5 show the cutaway view of HTTR reactor building, the schematic view of reactor components and reactor core structures and the RPV, and the cooling system of the HTTR, respectively [11,12]. Table 2.1 Major specification of HTTR [11]. Thermal power (MW)

30 

Outlet coolant temperature ( C) 

850/950

Inlet coolant temperature ( C)

395

Fuel

Low-enriched UO2

Fuel element type

Prismatic block

Direction of coolant flow

Downward flow

Pressure vessel

Steel

Number of main cooling loop

1

Heat removal

Intermediate heat exchanger (IHX) Pressurized water cooler (PWC)

Primary coolant pressure (MPa)

4

Containment type

Steel containment

Plant lifetime

About 20 years

24

High Temperature Gas-cooled Reactors

Figure 2.3 Cutaway view of HTTR reactor building [11].

Figure 2.4 Schematic view of components and reactor core vessel of HTTR [11].

Design of High Temperature Engineering Test Reactor (HTTR)

25

Figure 2.5 Cooling system of HTTR [12].

2.1.3.1 Reactor core The reactor core is composed of core components, reactor internals, reactivity control system, and the RPV as shown in Fig. 2.3.

2.1.3.1.1 Core components The reactor core consists mainly of hexagonal fuel blocks, control rod guide blocks, and replaceable reflector blocks. The active core, which is 2.9 m in height and 2.3 m in diameter, consists of 30 fuel columns and 7 control rod guide columns. One column is made up of five fuel blocks and four replaceable reflector blocks above and below the fuel blocks. The active core is surrounded by replaceable reflector blocks and permanent reflector blocks. The permanent reflector blocks are fixed tightly by the core restraint mechanism. A fuel assembly consists of fuel rods and a hexagonal graphite block, 360 mm across flats and 580 mm in height, as shown in Fig. 2.6. The fuel assembly has three dowels on the top and three mating sockets at the bottom to align the fuel assemblies. TRISO-CFPs with UO2 kernel, about 6 wt.% average enrichment and 600 μm in diameter, are dispersed in the graphite matrix and sintered to form a fuel compact. Fuel compacts are contained in a fuel rod, 34 mm in outer diameter and 577 mm in length. Fuel rods are inserted into vertical holes in the graphite block. Helium gas coolant flows through gaps between the holes and the rods.

26

High Temperature Gas-cooled Reactors

Figure 2.6 Fuel assembly of HTTR [11].

2.1.3.1.2 Reactor internals The reactor internals consist of graphite core support structures, metallic core support structures, and the other components as shown in Fig. 2.7. The graphite support structures consist of hot plenum blocks, core bottom structures (CBSs), core support posts, etc. The hot plenum blocks provide lateral and vertical positioning and support of the core array. The blocks contain flow paths, which guide the primary coolant from the outlet of the fuel columns and distribute it into the hot plenum beneath the hot plenum blocks. The core support posts are designed so as to support the core and hot plenum block arrays, which form the hot plenum. The permanent reflector block is a graphite structure surrounding the replaceable reflector blocks and control rod guide blocks located in the circumference of the active core. The metallic core support structures are composed of core support plates, a core support grid, and the core restraint mechanisms. The core support plate and the core support grid are placed below the thermal insulation layers. The core restraint mechanism surrounds the permanent reflector blocks.

2.1.3.1.3 Reactivity control system The control rods are individually supported by control rod drive mechanisms located in stand pipes connected to the hemispherical top head of the reactor vessel. The control rods are inserted into the channels in the active core and replaceable reflector regions around the active core. The control rod drive mechanism withdraws and inserts a pair of the control rods. Reactor shutdown is made at first by inserting nine pairs of control rods into the reflector region, and then by inserting the other seven pairs of the control rods into the active core region 40 min later or

Design of High Temperature Engineering Test Reactor (HTTR)

27

Figure 2.7 Reactor core internals of HTTR [1].

after the outlet coolant temperature is down to 750 C so that the control rods should not exceed their design temperature limit. Reserve shutdown capability is provided by insertion of B4C/C pellets into holes in the control rod guide blocks.

2.1.3.1.4 Reactor pressure vessel The RPV, 13.2 m in height and 5.5 m in diameter, is made of 21/4Cr1Mo steel. The 21/4Cr1Mo steel has better creep strength at high temperature than MnMo steel, which is widely used in the pressure vessels of light water reactors. The top head of the RPV is bolted to the flange of the cylindrical shell. Thirty-one stand pipes, including the control rod and the irradiation stand pipes, are welded to the top head. A stand pipe closure is installed on the top of each stand pipe and is removed during refueling.

2.1.3.2 Reactor cooling system The main cooling system of the HTTR is composed of a primary cooling system, a secondary helium cooling system, and a pressurized water-cooling system as schematically shown in Fig. 2.5. The primary cooling system, which has gas circulators and two heat exchangers, that is, a heliumhelium IHX and a PPWC, removes the heat from the reactor core to the secondary helium cooling system and pressurized water-cooling system. Primary helium gas is transferred from the core to the heliumhelium IHX and the PPWC through a primary concentric hot gas duct. The

28

High Temperature Gas-cooled Reactors

secondary helium cooling system, composed of the secondary pressurized water cooler (SPWC) and a gas circulator, removes the heat from the primary helium gas through the heliumhelium IHX. The pressurized water-cooling system consists of an air cooler and water pumps. The air cooler cools the pressurized water for both the PPWC and the SPWC, and transfers the heat from the reactor core to the final heat sink, that is, the atmosphere.

2.1.3.3 Engineered safety systems 2.1.3.3.1 Auxiliary cooling system The auxiliary cooling system consists mainly of the auxiliary heat exchanger, auxiliary gas circulators, and an air cooler as shown in Fig. 2.5. The auxiliary cooling system has a heat transfer capacity of about 3.5 MW. The auxiliary cooling system automatically starts up when the reactor is scrammed and the main cooling system is stopped by accidents. Core cooling by a forced circulation is possible with the auxiliary cooling system. The auxiliary cooling system consists of redundant dynamic components such as gas circulators, water pumps, and valves that are also operated with emergency power supply.

2.1.3.3.2 Vessel cooling system The VCSs consist of upper, lower, and side cooling panels, around the RPV, and cooling water circulation systems. The VCSs are used as a residual heat removal system when the forced circulation in the primary cooling system cannot be maintained due to the rupture of the inner pipe or both pipes in the concentric hot gas duct. The VCSs are also an engineered safety feature composed of two independent complete sets that are backed up with emergency power supply.

2.1.3.3.3 Containment structure The containment structure consists of a reactor containment vessel, service area, and an emergency air purification system, which reduce the release of FPs to the environment during the postulated accidents. The containment vessel is designed to withstand the temperature and pressure transients and to be leak-tight within the specified limits in the case of primary gas duct rupture. The service area is the space surrounding the containment vessel, where the fuel handling and storage systems and the primary and the secondary helium purification systems are located. The emergency air purification system removes airborne radioactivity and maintains pressure slightly lower than that of the atmosphere in the service area during the accidents.

2.1.3.4 Instrumentation and control system 2.1.3.4.1 Instrumentation system The nuclear instrumentation system of the HTTR is composed of a wide range monitoring system and a power range monitoring system. The wide range monitoring system is available as a postaccident monitor under accident conditions such as rupture of the primary concentric hot gas duct. The wide range monitoring system is used to measure the neutron flux from 1028% to 30% of rated power. Three

Design of High Temperature Engineering Test Reactor (HTTR)

29

fission chambers are installed in the permanent reflector blocks through the stand pipes. The power range monitoring system is used to measure the neutron flux from 0.1% to 120% of rated power. The power range monitoring system is also used as the sensor for the reactor power control system. The detectors of the power range monitoring system are located outside the RPV. Therefore the detectors have high sensitivities so as to detect the neutron flux at a very low level. In order to monitor the core outlet temperature of the primary coolant, seven thermocouples are arranged in the hot plenum blocks below the reactor core. N-type thermocouples (NicrosilNisil) are selected because the temperature of the primary coolant in the hot plenum reaches about 1000 C at rated power operation. The fuel failure detection system is composed of precipitators, a preamp, and a control box. The precipitator is used to detect β-rays radiated from short-lived gaseous FPs such as 88Kr, 89 Kr, and 138Xe.

2.1.3.4.2 Control system The reactor control system is designed to assure high stability and reasonably damped characteristics against various disturbances during operation. The main control system of the HTTR consists of the operational mode selector, the reactor power control, and plant control systems. The operational mode selector is designed to select several mode operations such as the rated power operation, the high temperature test operation, the safety demonstration test operation, and the irradiation test. The reactor power control system consists of the power control and reactor outlet coolant temperature control devices. The reactor outlet coolant temperature control device gives demand of reactor power to the power control device so that the outlet coolant temperature can be controlled to a certain value. The plant control system controls the plant parameters such as the reactor inlet coolant temperature, the primary coolant pressure, and differential pressure between the primary cooling system and the pressurized water-cooling system or secondary helium cooling system.

2.1.3.4.3 Safety protection system The safety protection system consists of the reactor protection and engineered safety features actuating systems. It is designed with a 2-out-of-3-circuit logic and 2-trains. The reactor protection system of the HTTR automatically initiates a reactor scram by inserting the control rods and simultaneously stops gas circulators. The engineered safety features actuating system of the HTTR is designed to arrest the release of the FP and to ensure the integrity of the core, the reactor coolant pressure boundary, and the containment vessel boundary against unexpected conditions during abnormal operational transients and accidents such as a primary cooling system pipe rupture.

2.1.4 R&D programs for HTTR 2.1.4.1 Fuel Fabrication technology of high-quality CFPs with low failure fraction and high irradiation resistance was developed. Irradiation tests on the fuel performance under normal operating conditions were conducted using the Oarai Gas Loop No. 1

30

High Temperature Gas-cooled Reactors

(OGL-1), the gas-swept capsules, and the closed capsules at the Japan materials testing reactor (JMTR), and the closed capsules at the Japan research reactor-2 (JRR-2). In parallel, the fuel behavior under accident conditions was also investigated in out-of-pile ramped and isothermal-heating tests on irradiated CFPs.

2.1.4.2 Graphite High-grade graphite IG-110 with high strength, corrosion, and irradiation resistance was developed and graphite structural design guideline was established. The studies on mechanical properties and nondestructive test of the nuclear grade graphite were performed to support the design and acceptance inspection for the HTTR core and core support components.

2.1.4.3 Metallic materials Heat and corrosion-resistant superalloy Hastelloy XR was developed, which can be used at temperatures as high as 950 C at normal operation and 1000 C in accidents, and high temperature structural design guideline for high temperature metallic components was established. Comprehensive qualification tests such as creep, fatigue, and corrosion on Hastelloy XR were carried out to accumulate the test data for structural design and safety evaluation. On the other hand, R&D on a long-term target alloy, NiCrW superalloy, was carried out for application at service temperatures around 1000 C.

2.1.4.4 Reactor physics Reactor physics experiments were conducted using a very high temperature reactor critical assembly (VHTRC). Seven different cores having radial and axial reflectors were assembled at the VHTRC to study the detailed neutronic characteristics of the HTTR core. The VHTRC-1 and -2 cores were loaded mainly with 4 wt.%-enriched fuel and the VHTRC-3 core was loaded mainly with 6 wt.%-enriched fuel. These three cores had an axially uniform loading pattern. On the other hand, VHTRC-4, -5, -6, and -7 cores were loaded with 2, 4, and 6 wt.%-enriched fuel rods in axially zoning patterns.

2.1.4.5 Reactor instrumentation In the field of nuclear instrumentation for reactor operation and control, hightemperature fission counter-chambers (HTFCs), gamma-compensated ionization chambers (CICs), and high-sensitive gamma-uncompensated ionization chambers (HSUICs) were developed. The HTFCs and the HSUICs are adopted to the wide range monitoring system and the power range monitoring system of the HTTR, respectively. As for in-core temperature sensors, the advanced NiCr thermocouples, that is, NicrosilNisil thermocouples (N-type TCs), were developed. The N-type TCs have a Nicrosil sheath with a heat-resistant ceramic coat of about

Design of High Temperature Engineering Test Reactor (HTTR)

31

200 μm in thickness. They showed stable electromotive force characteristics under out-pile tests at 1200 C for 20,000 h and in-pile irradiation tests at 1000 C.

2.1.4.6 Heat transfer and fluid dynamics 2.1.4.6.1 Air ingress process following primary-pipe rupture When the primary-pipe rupture accident occurs, the high-pressure helium gas coolant in the reactor is forced out into the reactor containment vessel through the breach. Gas pressure would be balanced between inside and outside of the RPV after a few minutes. After that, it is supposed that air enters the reactor core from the breach due to molecular diffusion and natural convection of a multicomponent gas mixture. In the first stage of the accident, density of the gas mixture in the reactor gradually increases as air enters by the molecular diffusion and natural convection of the gas mixture. Finally, natural circulation of the air occurs suddenly throughout the entire reactor and the second stage of the accident starts. The process of air ingress during the first and second stages was investigated experimentally and numerically. According to the test results, the rate control process of air ingress in the first stage is the molecular diffusion and the very weak natural circulation, which results in a small amount of air ingress rate. On the other hand, air flows into the reactor by the ordinary natural circulation in the second stage, which results in a large amount of air ingress rate.

2.1.4.6.2 Graphite oxidation in case of air ingress into reactor core Graphite corrosion rates were investigated experimentally as well as theoretically with a particular focus on corrosion by high temperature gas stream. A coefficient of the graphite corrosion rate was calculated and measured in wide range of gas temperature and O2 concentration. Experiments of the graphite oxidation at the mass transfer controlled regime (high temperature region) were performed during the natural circulation in the air ingress accidents. The mass transfer coefficients and the generation ratio of CO and CO2 were obtained.

2.1.4.7 Components and structures at high temperature The helium engineering demonstration loop (HENDEL) was constructed for performing full-scale demonstration tests on the core internals and high temperature components for the HTTR. In the fuel stack test section (T1) of the HENDEL, thermal and hydraulic performances of helium gas flowing through a fuel rod channel and a fuel stack were investigated for the HTTR core thermal design. In addition, functioning reliability of a control rod drive mechanism and a control rod assembly was confirmed using a mockup model. On the other hand, demonstration tests were conducted to verify thermal and hydraulic characteristics and structural integrity related to the CBS using a full-scale test facility named as the in-core structure test section (T2). For example, sealing performance tests revealed that leakage of low temperature helium gas through gaps between the permanent reflector blocks to the core is at very low level compared with the HTTR design value and no change of the leakage flow rate was observed after a long-term operation.

32

High Temperature Gas-cooled Reactors

2.2

Nuclear design

2.2.1 Introduction The HTTR is a block-type HTGR, designed for 950 C outlet gas temperature. The high outlet gas temperature may increase the FP release from fuel because of increased fuel temperature. Therefore it was important in the nuclear design to suppress the fuel temperature rise by optimizing the power distribution, which was carried out in the following steps. First, we planned to replace each fuel block with a fresh fuel block after every burnup cycle. The axial and radial power distribution was optimized to flatten by allocating optimally the fuel blocks with 12 different uranium enrichments of fresh fuel throughout the core. Second, the loading of fresh fuel blocks throughout the core caused large initial excess reactivity, which resulted in the deep insertion of control rods into the core and resulted in high fuel temperatures. Greater excess reactivity was necessary because fast reactivity depletion resulted from the low conversion ratios of the lowenriched uranium fuel. Thus it was important for the nuclear design of the HTTR to reduce the excess reactivity adequately. The problem with the excess reactivity was solved by optimizing burnable poisons (BPs) in the core. The optimization kept the excess reactivity to the minimum necessary for reactor operations. Third, deviation from the optimum power distribution due to burnup of fissile materials was avoided by optimizing the specifications of BPs, namely, the poison atom density and the radius for each local area. The excess reactivity of the core was maintained constant. It becomes possible, then, to operate the reactor without changing the insertion position of control rods during power operation at 950 C. Based on the above design policy, the design requirements, analytical methods, and nuclear characteristics of the HTTR were determined, and the analytical methods were improved based on experimental results [13,14].

2.2.2 Design requirement The main physical parameters of the HTTR core are listed in Table 2.2. The nuclear design must satisfy the following requirements.

2.2.2.1 Excess reactivity Excess reactivity must be determined taking into account the following effects: G

G

G

G

G

temperature increase from the cold shutdown state to the rated power operation state; build-up of FPs, such as 135Xe and 149Sm; burnup of fuel; margins including the irradiation tests; and uncertainty for nuclear calculations.

2.2.2.2 Reactor shutdown margin The control rods must be so designed as to provide a reactor shutdown margin of more than 0.01 Δk/k, even if one pair of control rods having the maximum

Design of High Temperature Engineering Test Reactor (HTTR)

33

Table 2.2 Specifications of HTTR [14]. Thermal power (MW)

30

Outlet coolant temperature ( C)

950

Inlet coolant temperature ( C)

395

Primary coolant pressure (MPa)

4

Core structure

Graphite 2.3

Equivalent core diameter (m)

2.9

Effective core height (m) 3

2.5

Average power density (W/cm ) Fuel

UO2

Uranium enrichment (wt.%)

310

Type of fuel

Pin-in-block

Burnup period (days)

660

Fuel compact Outer diameter (cm)

2.6

Inner diameter (cm)

1.0

Length (cm)

3.9

Packing fraction of CFPs (vol.%) 3

Density of graphite matrix (g/cm ) Impurity in graphite matrix (ppm)

30 1.7 ,1.2 (boron equivalent)

CFP CFP diameter (μm)

920

Kernel diameter (μm)

600

3

Density (g/cm )

10.41

Coating material

PyC/PyC/SiC/PyC

Layer thickness (μm)

60/30/25/45

Coolant Material

Helium gas

Flow in core

Downward

Reflector thickness Top (m)

1.16

Side (m)

0.99

Bottom (m)

1.16

Number of fuel blocks

150

Number of fuel columns

30

Number of pairs of control rods

16

reactivity worth is completely withdrawn and cannot be reinserted. The reserved shutdown system must be designed to allow for a reactor shutdown margin of 0.01 Δk/k or more, even if the control rod system is not available.

34

High Temperature Gas-cooled Reactors

2.2.2.3 Reactivity addition rate The maximum reactivity addition rate with the control rods must be limited to such extent that a related power excursion does not impair the integrity of the core, the reactor internal structures, the primary cooling system, etc. In order to satisfy these conditions, the control rod system shall be so designed that the withdrawal length in one step is limited below 50 mm and the maximum reactivity addition rate does not exceed 2.4 3 1024 Δk/k/s, even if the control rods are withdrawn with the permissible maximum speed.

2.2.2.4 Reactivity coefficient An important reactivity coefficient is the power coefficient of reactivity, which is dominated by the Doppler and the moderator temperature coefficients. The reactor core must have a negative reactivity feedback characteristic, which dampens the power level change. To achieve this condition, the reactor core must be designed in such a way that the power coefficient of reactivity is negative for any operation condition.

2.2.2.5 Power distribution The power distribution must be so determined that the fuel temperature does not exceed the limited value during operation. To obtain this condition, the fuel and BP must be loaded so that the maximum fuel temperatures are kept as low as possible.

2.2.2.6 Burnup The maximum average burnup in a fuel element must not exceed 33,000 MWd/t. To achieve the outlet coolant temperature of 950 C, 235U enrichment distribution and BP distribution in the core are optimized. Therefore burnup of fuel becomes low.

2.2.3 Analytical method The shutdown margin, control rod worth, reactivity coefficient, and power distribution must satisfy the design requirements. Therefore they should be evaluated with a nuclear design code system (NDCS) of sufficient accuracy. The calculation accuracy of the system was evaluated from the measured values of the VHTRC.

2.2.3.1 Design codes The NDCS for the HTTR [15] consists of the computer codes DELIGHT, TWOTRAN2 [16], and CITATION-1000VP [17]. The program structure of the system is shown in Fig. 2.8. DELIGHT is a one-dimensional lattice burnup cell calculation code, which has been developed in JAERI especially for the nuclear design of the HTTR. This code is used to provide group constants of fuel blocks and graphite blocks for the succeeding core calculation. TWOTRAN-2 is a transport code, which is used to provide the average group constants of a graphite block where control rods are inserted. CITATION-1000VP is a reactor core analysis code, which has been improved from CITATION in JAERI to perform the nuclear characteristics analysis with a whole core model of the HTTR.

Design of High Temperature Engineering Test Reactor (HTTR)

35

Specifications of fuel, burnable poison, graphite blocks, control rod, core, reflector, etc. DELIGHT Calculating neutron spectrum and providing the group constants of fuel blocks and graphite blocks

TWOTRAN-2 Providing the group constants of control rod-inserted graphite blocks

Group constants set CITATION-1000VP Analysis of reactor core characteristics 1. Effective multiplication factor 2. Neutron flux and power distribution 3. BP rod worth 4. Control rod worth, etc. Figure 2.8 Structure of the nuclear design code system for HTTR [14].

2.2.3.2 Validation of design code using very high temperature reactor critical assembly The critical assembly VHTRC was constructed to obtain data for validation of the NDCS. Calculation errors of effective multiplication factor, neutron flux distribution, BP reactivity worth, control rod worth, and temperature coefficients by the NDCS were confirmed for the NDCS. Calculation errors of Monte Carlo code were also evaluated. The Monte Carlo code showed good agreement with experimental data of the VHTRC [18,19]. The Monte Carlo code could also be applied for the evaluation of HTGR’s design.

2.2.4 Evaluation of nuclear characteristics 2.2.4.1 Excess reactivity and nuclear shutdown margin The reactivity of the core is controlled by the control rods, BP rods, and reserved shutdown system. The control rods are used for power regulation and shutdown during normal operation because the primary coolant flow rate is kept constant during operation. The BP rods are used to obtain an adequate initial excess reactivity for the operation. The locations of the control rods, BP rods, and reserved shutdown system in the core are shown in Fig. 2.9. The excess reactivity, reactor shutdown margin and functions of the control rods, BP rods, and reserved shutdown system are described in the following.

2.2.4.1.1 Excess reactivity The excess reactivity was determined to compensate for reactivity losses. The details of the reactivity losses and excess reactivity are as follows.

36

High Temperature Gas-cooled Reactors

: Fuel column N : Zone number : Burnable poison : Control rod guide column

*1 235U enrichment (wt%) *2 Natural boron concentration (wt%) *3 The number indicates the layer number froom the top fuel block

: Control rod : Reserve shutdown system : Replaceable reflector : Irradiation column

Figure 2.9 Fuel and BP configuration plan of HTTR core [14].

Reactivity losses 1. 2. 3. 4. 5.

Reactivity loss from the cold shutdown state to the rated power operation state. Build-up of rapidly saturating FPs, such as 135Xe and 149Sm. Depletion of fissile material and the build-up of heavy metal isotopes. Build-up of slowly saturating FPs. Reactivity margin of irradiation tests.

The reactivity losses of (1) and (2) are about 0.064 Δk/k and 0.024 Δk/k, respectively. The sum of the reactivity losses of (3) and (4) is about 0.043 Δk/k. The total of the reactivity losses is about 0.131 Δk/k. Excess reactivity The excess reactivity is determined to compensate for the reactivity losses and provide a reactivity margin for operation. The reactivity margin for operation is about 0.003 Δk/k. The excess reactivity is then about 0.165 Δk/k in consideration of calculation uncertainties (0.016 Δk/k) and reactivity increase due to the irradiation test of fuel (0.015 Δk/k).

2.2.4.1.2 Burnable poison rods If the core does not contain the BP rods, the excess reactivity of the core becomes too high to control the reactivity by the control rods only. The excess reactivity is reduced by the neutron absorption effect of the BP rods to an adequate value as

Design of High Temperature Engineering Test Reactor (HTTR)

37

Figure 2.10 Change in excess reactivity in rated power operation state where control rods are fully withdraw [14].

shown in Fig. 2.10. In cold condition, the reactivity compensated by the BP rods is 0.113 Δk/k at beginning of life (BOL).

2.2.4.1.3 Controllable reactivity and shutdown margin Control rod Seven pairs of the control rods are inserted into the active core, and nine pairs into the removable reflector region. The control rods are designed so that the controllable reactivity is larger than the excess reactivity to achieve a sufficient shutdown margin. The controllable reactivity of the control rods corresponds to the core reactivity, which is held by full insertion of all control rods except for two pairs. One is the pair at the core center. This pair is dismantled during irradiation tests. The other is a pair of the control rods neighboring the core center. This pair has the maximum reactivity worth among all pairs of the control rods, and it should be considered for the shutdown margin evaluation to be inoperable at the fully withdrawn position. The controllable reactivity, in this case, is at least 0.18 Δk/k taking into the calculation uncertainties, and the shutdown margin more than 0.015 Δk/k. Reserved shutdown system The absorbers of the reserved shutdown system are pellets produced by sintering B4C powder and graphite. These pellets can be released from the storage hoppers into the reserved shutdown system holes in the control rod guide columns to act as a backup system for the control rod system. The reserved shutdown system in the core center is not involved in measuring the controllable reactivity worth since the control rod guide column is not used for the reserved shutdown

38

High Temperature Gas-cooled Reactors

system during the irradiation test. The controllable reactivity of the reserved shutdown system is at least 0.18 Δk/k for any operating control rod configuration including consideration of calculation uncertainty. The reserved shutdown system must be capable of compensating the reactivity, which arises in the change from the full-power operation state to the cold shutdown state. This reactivity increase is less than 0.088 Δk/k for any operation condition through a burnup cycle. The shutdown margin of the reserved shutdown system is about 0.092 Δk/k for all operating conditions.

2.2.4.2 Reactivity addition rate and reactivity coefficient 2.2.4.2.1 Reactivity addition rate The reactivity addition rate results from the driving speed and reactivity worth of a control rod. The maximum driving speed of the control rod is 60 cm/min. The maximum control rod reactivity worth per unit length is 2.3 3 1024 Δk/k/cm. The maximum reactivity addition rate is then 2.3 3 1024 Δk/k/s.

2.2.4.2.2 Reactivity coefficient Reactivity coefficients include the Doppler, moderator temperature, and power coefficients. The Doppler coefficient is the rate of reactivity change due to a change in the average fuel temperature, and is negative for any operating conditions as shown in Fig. 2.11. The moderator temperature coefficient is the rate of reactivity change due to a change in the moderator temperature. It is largely negative for all temperatures at BOL, but could become positive within a certain temperature range at end of life (EOL), as shown in Fig. 2.12. The power coefficient is the rate of reactivity change due to a

Figure 2.11 Doppler coefficients at BOL and EOL [14].

Design of High Temperature Engineering Test Reactor (HTTR)

39

Figure 2.12 Moderator temperature coefficients at BOL and EOL [14].

change in the reactor power level. The power coefficient, dominated by the combination of the Doppler and the moderator temperature coefficients, remains negative at any temperature even if there is the possibility of positive moderator temperature coefficient.

2.2.4.3 Power distribution and burnup 2.2.4.3.1 Power distribution

To achieve the high reactor outlet coolant temperature of 950 C, the 235U enrichment and natural boron concentration of the BP rod in each fuel block are determined to keep the maximum fuel temperature as low as possible. The 235U enrichment and natural boron concentration are also shown in Fig. 2.9. To keep the radial power distribution unchanged through the burnup cycle, the 235U enrichment in the outer region has to be larger than in the inner region. To create an axial fuel temperature distribution as uniform as possible, the 235U enrichment in the upper region has to be larger than that in the lower region. This is due to the large temperature difference of 550 C between the inlet and outlet of coolant. Radial power distribution The radial power peaking factors are estimated from the coarse power distribution for various burnup steps. Table 2.3 shows that the radial power peaking factor does not exceed 1.1. Thus the radial power distribution appears to be sufficiently flattened through all of the burnup cycles in uniform fuel temperature distribution. As for the effect of heterogeneity caused by the fuel rods, BP rods, etc., the correction factor of local radial power peaking is estimated and

40

High Temperature Gas-cooled Reactors

Table 2.3 Radial power peaking factor [14]. Burnup (days)

Fuel zone 1

Fuel zone 2

Fuel zone 3

Fuel zone 4

0

1.04

1.07

1.08

1.10

10

1.06

1.08

1.09

1.10

110

1.08

1.08

1.08

1.10

220

1.09

1.08

1.09

1.10

330

1.09

1.08

1.08

1.09

440

1.10

1.08

1.08

1.09

550

1.10

1.08

1.08

1.10

660

1.09

1.09

1.09

1.10

Figure 2.13 Moderator temperature coefficients at BOL and EOL [14].

confirmed to be less than 7%. This correction factor is accounted for in the maximum fuel temperature evaluation. Axial power distribution The axial power distribution for several burnup steps is shown in Fig. 2.13. At BOL, the power is larger in the upper region because the 235U enrichment is higher than that of the lower region. With burnup progress, the power in the upper region decreases considerably 440 days after BOL. Then the control rods are withdrawn gradually, and the power in the upper region increases again. In spite of these power changes, the distribution of the fuel temperature in the core remains very uniform. As for the effect of heterogeneity caused by the fuel and reflector block structure, the correction factor of local axial power peaking is estimated and confirmed to be below 4%. This correction factor is also accounted for in the maximum fuel temperature evaluation.

2.2.4.3.2 Burnup The average burnup of each fuel block at EOL is shown in Table 2.4. The maximum burnup appears in the second layer in the second fuel zone, and amounts to

Design of High Temperature Engineering Test Reactor (HTTR)

41

Table 2.4 Average burnup (MWd/t) of each fuel block at EOL [14]. Axial fuel layer

Fuel zone 1

Fuel zone 2

Fuel zone 3

Fuel zone 4

1

20,500

21,000

20,500

21,000

2

31,000

31,500

30,500

30,500

3

27,500

28,500

27,000

26,500

4

18,500

19,000

17,500

17,500

5

13,000

13,000

12,000

12,000

31,500 MWd/t. Thus even considering the uncertainty in the heavy metal inventory, the maximum burnup remains within the design limit of 33,000 MWd/t. The average burnup of the core is about 22,000 MWd/t. To achieve 950 C of outlet coolant temperature, 235U enrichment distribution and specification of BP are optimized. Therefore the burnup became relatively low.

2.3

Core thermal-hydraulics

2.3.1 Introduction The HTTR is a helium-cooled, graphite-moderated high temperature reactor with 30 MW thermal power and a reactor outlet coolant temperature of 950 C. This maximum outlet coolant temperature of 950 C is achieved in the high temperature test operation mode. The core thermal-hydraulic design is carried out considering specific characteristics of the HTTR such as a prismatic fuel element and coated fuel particles, the nuclear and core component design, hot-spot factors, fuel design limit, etc. Approximately 95% of the 30 MW thermal power is generated in the fuel elements and the rest in the graphite moderator. Thermal and irradiation conditions become severe for the graphite blocks at the outermost region of the core so that fuel elements in the region contain only 31 fuel rods instead of 33 to keep their structural integrity. The coolant, flowing downward in the annular space between a fuel rod and a hole in the graphite block, is heated to about 1000 C at the exit of the core. Approximately 97% of the reactor thermal power is removed by the main cooling system with a helium flow of 12.4 kg/s at rated operation and 10.2 kg/s at the high temperature test operation. The rest of the heat is removed, mainly, by the VCS. After reactor scram, the residual heat of the core is removed by the auxiliary cooling system and/or the VCS. This section describes outline of the core thermal-hydraulic design of the core and the analytical estimation of the maximum fuel temperature based on the experimental results of the HTTR [20].

2.3.2 Design requirements As for the integrity of the fuel itself, it is necessary to maintain the maximum fuel temperature as low as possible under normal operation and any AOOs. The fuel

42

High Temperature Gas-cooled Reactors

design limit is determined considering the experimental results of in-pile as well as out-of-pile fuel performance tests simulating reactor operation. The failure of CFPs is caused mainly by kernel migration (amoeba effect) and corrosion of the SiC layer by palladium at high temperatures and a high temperature gradient in the fuel compact. Major safety requirements for the fuel to be maintained intact, especially, in respect of the core thermal-hydraulic design, are as follows: G

G

The CFPs shall not fail significantly during normal operation. The maximum fuel temperature shall not exceed fuel design limit of 1600 C during any AOOs. It is confirmed experimentally that the coating layers of the CFPs would maintain their intactness below 1600 C.

To satisfy these requirements, the maximum fuel temperature limit for normal operation is specified as 1495 C.

2.3.3 Design details Fig. 2.14 shows a calculation flow diagram of the core design including the nuclear and thermal-hydraulic design. The coolant flow rate and coolant temperature distributions are evaluated by the flow network analysis code FLOWNET. The

Figure 2.14 Calculation flow of core design [20].

Design of High Temperature Engineering Test Reactor (HTTR)

43

Figure 2.15 FLOWNET analysis model for HTTR: (A) axial flow network model, (B) radial flow network model [20].

calculation model of FLOWNET consists of one-dimensional flow branches and pressure nodes, which are the junctions or terminals of the branches as shown in Fig. 2.15. Each branch is provided with an equivalent cross-section area, length, hydraulic diameter, and pressure loss coefficient of the actual passage. Pressure loss coefficients used in the calculation for the leakage flow between permanent reflectors and/or hot plenum blocks and the cross-flow between fuel elements, control rod guide blocks, and replaceable reflector blocks were derived from experimental results. A detailed coolant flow distribution is determined based on the power distribution obtained in the nuclear design and the dimensions of the core components and reactor internals. Flow paths, regarding the coolant flow analysis, are the main coolant flow in the graphite blocks, the bypass flow in the intercolumn gaps, the leakage flow through the permanent reflectors and the cross-flow in the horizontal interface gaps of the hexagonal graphite blocks. To achieve a high reactor outlet coolant temperature of 950 C, it is important to keep the maximum fuel temperature as low as possible. Therefore in the core thermal-hydraulic design, the coolant flow, ineffective to direct fuel cooling, should be minimized. Since the reduction of the ineffective coolant flow such as leakage, cross and bypass flows cannot be eliminated completely in the design of a prismatic fuel core, the plan is to seal the core bypass flow at the outer part of the core. Pressure loss coefficients for each flow branch and the sealing mechanism are derived from experimental results. The fuel temperature is calculated by the fuel temperature analysis code TEMDIM, using a cylindrical model, based on the power distribution including local power peaking, coolant flow distribution including redistribution in the fuel column, and hot-spot factors. The hot-spot factors or hot channel factors are considered in the core thermalhydraulic design to evaluate the maximum fuel temperature not only during normal operation condition but also during any AOOs with an adequate conservativeness. They are also used to account for various uncertainties and to assure that the specified maximum fuel temperature in the core does not exceed the fuel design limit at any time and any location during the normal operation condition. Hot-spot factors consist of systematic factors (direct accumulation of conservatism) such as total reactor

44

High Temperature Gas-cooled Reactors

power, coolant flow rate, and inlet coolant temperature, and random factors (statistically treated), such as manufacturing tolerances and uncertainties on physical properties.

2.3.4 Evaluation results of design As a result of the coolant flow analysis, an effective core flow rate, as high as 88% of the total flow rate, is achieved for the minimum value. The effective core flow rate is calculated with the worst combination of flow passages and pressure loss coefficients. The coolant flow rate in the fuel-cooling channel decreases in the flow direction due to the cross-flow. The minimum coolant flow rate occurs at the fourth layer of the fuel. The cross-flow in the upper part of the core flows out of the fuel-cooling channel; in the lower part it flows in the opposite direction. The total leakage flow through the gaps between the permanent reflector blocks is approximately 1.1% of the total coolant flow. Moreover, the effects of deviations of the coolant flow rate from nominal value, which are caused by column bending, distribution of intercolumn gap width, manufacturing tolerances, etc., are estimated by the sensitivity analysis to determine hot-spot factors for the fuel temperature analysis. Approximately 3.0% of the coolant decreases from the nominal flow rate due to column bending, and is considered as one of the systematic factors relating coolant flow rate. The value of 3.2% is determined as the random factor due to the intercolumn gap width distribution and manufacturing tolerances. The maximum fuel temperature corresponds to the inner surface temperature of the fuel compact and is shown in Fig. 2.16. The solid lines in the figure show nominal values and the dotted line, the maximum fuel temperature considering hot-spot factors.

Figure 2.16 Illustration of axial temperature distribution for 950 C operation [20].

Design of High Temperature Engineering Test Reactor (HTTR)

45

Figure 2.17 Radial fuel temperature distribution at 950 C operation [20].

The maximum fuel temperature is achieved during the high temperature test operation, and is 1492 C after approximately 440 effective full-power operation days. It does not exceed the 1495 C of the fuel design limit for normal operation. Fig. 2.17 shows the radial temperature distribution at the high temperature test operation. On the other hand, the maximum fuel temperatures for the rated and the irradiation test operation are 1420 C and 1456 C, respectively, and are less than the temperature for the high temperature test operation. It has also been confirmed that the maximum fuel temperature during an AOOs is 1555 C for a case of abnormal control rod withdrawal at rated operation. In this case, the reactor does not be scrammed because reactivity insertion is below cents.

2.3.5 Reevaluation of maximum fuel temperature with operational data 2.3.5.1 Revision of calculation condition In the thermal-hydraulic design, the maximum fuel temperature is evaluated by using the power distribution calculated by core physics codes, and the coolant flow rate in the fuel channel, calculated by the FLOWNET code, and taking account of hot-spot factors as shown in Section 2.3.4. To reconfirm the fuel integrity during the hightemperature test (950 C) operation mode, the fuel temperature was re-evaluated taking account of the measured data in the rise-to-power test of 850 C. The reevaluation of the maximum fuel temperature was performed with the same method as in the thermal-hydraulic design and the revised hot-spot factors by measurement data through rise-to-power test, gamma ray measurement of fuel block. As for the power distribution and coolant flow rate in the fuel channel, it was confirmed that the core calculation code and FLOWNET code gave valid calculation result as follows. Power distribution in each fuel block was evaluated by gamma ray from each fuel block. Comparison of measured and calculated results of power distribution in

46

High Temperature Gas-cooled Reactors

Figure 2.18 Comparison of experimental and analytical results for power density [20].

Figure 2.19 Comparison of experimental and analytical results for coolant temperature rising at core [20].

each fuel column at 9 MW is shown in Fig. 2.18. There is no significant difference between measured and calculated results. Considering uncertainties of measurements of power distribution, calculated power distribution shows good agreement with measured results. Comparison of measured with calculated results of coolant temperature, in the hot plenum block, was carried out. The coolant temperatures in the hot plenum block were calculated by FLOWNET. Comparison of coolant temperatures, in the hot plenum block, is shown in Fig. 2.19. Measured results show higher values at the center. The calculated results also show higher values at the center and show similar temperature differences between the center and the outer region to the measured results. It is concluded that the flow distribution in the HTTR core calculated by FLOWNET code is reliable. Hot-spot factors, such as reactor power estimation error and power distribution error, were revised based on measured data. For example, a hot-spot factor of uncertainty of reactor thermal power consisted in 2% of measurement error and 0.5% of control systems in the thermal-hydraulic design. In the reevaluation, this hot-spot factor was revised as 1.5% based on an error of measurement system and operation data of HTTR. Moreover, calculation conditions such as core inlet

Design of High Temperature Engineering Test Reactor (HTTR)

47

temperature, total core flow rate, and control rod positions were revised by operation data of the HTTR for realistic evaluations.

2.3.5.2 Reevaluation result by operational data As the result of reevaluation, maximum fuel temperature was estimated as 1463 C at the high temperature test operation at 160 days of effective full-power operation days. This result is below the limit temperature for normal operation of 1495 C; it is also below the results of the thermal-hydraulic design of the HTTR. An effective core flow rate becomes 90% of the total flow rate. Compared to the results of the thermal-hydraulic design, ratio of coolant flow in a fuel-cooling channel is larger than that in the thermal-hydraulic design. It is due to the revision of calculation conditions of power distribution and flow rate distribution according to operational data. Axial temperature distributions of coolant and fuel in a fuel column show maximum fuel temperature are depicted in Fig. 2.20. While maximum fuel temperature occurs at the bottom of fuel column in the thermal-hydraulic design, as shown in Fig. 2.16, maximum fuel temperature occurs at middle of fuel column in the reevaluation. It is due to a change in axial power distribution. In the reevaluation of power distribution, the calculation conditions of control rod position are higher than that in the thermal-hydraulic design. The control rod position is found according to the experimental data. An axial power distribution that shows a “sign curve” is affected by the control rod position. The control rod position becomes higher; thus the position of the axial power peak becomes higher in the core. Therefore the position of the maximum fuel temperature moves to higher positions. Maximum fuel temperature distribution in each fuel column is shown in Fig. 2.21. In the thermal-hydraulic design, fuel column in the core center shows

Figure 2.20 Illustration of axial temperature distribution for 950 C operation [20].

48

High Temperature Gas-cooled Reactors

Figure 2.21 Radial fuel temperature distribution at 950 C operation [20].

higher fuel temperature. In the reevaluation, the outer fuel column shows higher fuel temperature. In the HTTR, radial power distribution is adjusted to be flat during burnup period by a fuel enrichment distribution in the core. However, the radial power distribution changes slightly with increase in burnup. In the thermalhydraulic design, a maximum fuel temperature of 1492 C appears at 440 days of burnup. The temperature distribution in Fig. 2.21 is the result at 160 days of burnup. The difference of position where maximum temperature occurs is due to the difference of radial power distribution for different burnup.

2.4

Graphite components

2.4.1 Introduction The HTTR is a graphite-moderated and helium gas-cooled reactor with the core composed of prismatic fuel elements of hexagonal graphite blocks. The characteristics of graphite are quite different in stressstrain response from metals, since the ductility of graphite is significantly less than the ductility of metals. Needless to say, design codes for metal components cannot be applied directly to the graphite components. JAEA had to develop the design criteria taking account of the brittle fracture behavior. The concept and key specification of the developed graphite design criteria are described, and also an outline of the quality control specified in the design criteria is mentioned.

2.4.2 In-core graphite and carbon structure in high temperature engineering test reactor The reactor core is an array of graphite blocks (fuel assemblies, control rod guide blocks, and replaceable reflector blocks), which provide the physical structure for arrangement and confinement of the fissile fuel materials, neutron moderation, heat

Design of High Temperature Engineering Test Reactor (HTTR)

49

transfer, and the positioning of control/shielding absorber materials. The core is supported by the core support structures and fixed by the core restraint mechanism [11]. As shown in Fig. 2.4, the graphite components are divided into two kinds, one for the permanently installed graphite components of the core support structures and the other for the replaceable components of the reactor core. IG-110 graphite, PGX graphite, and ASR-0RB carbon are used for the in-core components.

2.4.2.1 Core graphite components The fuel graphite blocks are graphite hexagonal right prisms with an array of fuel holes. The fuel graphite blocks are 360 mm in across flats and 580 mm height as shown in Fig. 2.6. Three dowel pins are installed on the top face, and they engage with dowel socket in the bottom face of the block above. The dowel arrangement ensures the correct orientation of fuel graphite blocks within the column with respect to each other. They are fabricated from IG-110 graphite, isotropic fine grained nuclear grade graphite. Table 2.5 shows typical thermomechanical properties of the IG-110 and PGX graphite, and ASR-0RB carbon [21]. Control rod guide blocks and replaceable reflector blocks have the same external shape as the fuel blocks. They are also fabricated from the IG-110 graphite.

2.4.2.2 Core support graphite components The hexagonal hot plenum block array is made up of two axial layers. This structure provides lateral and vertical positioning and support of the core array. The hot plenum block assembly contains passages, which collect the primary coolant flow from the outlet of the columns and distribute it into the high temperature plenum beneath the hot plenum blocks. Hot plenum blocks operate at core outlet gas

Table 2.5 Typical thermomechanical properties of graphite and carbon materials (unirradiated and unoxidized condition) [21]. IG-10

PGX

1.78 3 10

Mean tensile strength (MPa, 300 K)

25.3

8.1

6.8

Mean compressive strength (MPa, 300 K)

76.8

30.6

50.4

Young’s modulus (GPa, 300 K) ( 6 1/3Su)a

7.9

6.5

8.7

Mean thermal expansion coefficient (293673 K) (1026/K)

4.06

2.34

4.40

3

1.73 3 10

ASR-0RB

Bulk density (kg/m , 300 K)

3

3

1.65 3 103

Thermal conductivity (W/mK, 600 K)

80

75

10

Ash (ppm)

Max. 100

Max. 7000

Max. 5000

Grain size (μm)

Mean 20

Max. 800

Max. 2000

a

Determined from the cord joining two points (one point is the one-third of the specified minimum tensile strength and the other is the one-third of the specified minimum compressive strength) on the stressstrain curve.

50

High Temperature Gas-cooled Reactors

temperature, 950 C. These blocks are fabricated from PGX graphite, a structural grade, medium-to-fine grained molded graphite. The core support posts and seats are designed to structurally support the core and hot plenum block array while providing a flow plenum to receive the primary coolant flow exiting the core. The posts and seats are made of IG-110 graphite. The thermal insulation layer at the core bottom consists of three blocks: lower plenum block, carbon block, and bottom block. The main function of the carbon block, which is fabricated from nuclear grade ASR-0RB carbon, is to keep the metallic core support structures below 500 C. The permanent reflector is a graphite structure immediately surrounding the replaceable reflector and control rod guide columns located in the circumference of the core. The permanent reflector is an assembly of graphite blocks making 12 circumferential segments and 8 axial layers.

2.4.3 Concepts of graphite design criteria The design criteria were developed by partially modifying the ASME CE Code in the items of by-axes failure theory, buckling limit, oxidation effects on the basis of test data. The design criteria for the graphite components in the HTTR have been developed by JAEA [22]. The limits in the JAEA design criteria for HTTR are detailed in the following.

2.4.3.1 Component classification The graphite components have different use and functions as follows: 1. The core graphite components are replaced at regular intervals; on the other hand, the core support graphite components are permanent components, not replaced as shown in Table 2.6. 2. The core support components have more serious structural functions than the core components, that is, the core support components are thought to be more important from a safety viewpoint.

Therefore the graphite components are divided into two components, core components and core support components, and the stress limit for the core support graphite components is specified to be more severe than for the core graphite components. Table 2.6 Differences between core and core support graphite components [21].

Main component

Core graphite components

Core support graphite components

Fuel block

Hot plenum block

Graphite sleeve

Permanent reflector block

Control rod guide block

Core bottom structure

Replaceable reflector block

Core support post

Replaceability

Routine

Irradiation effects

Major

Design life

3 years

Design of High Temperature Engineering Test Reactor (HTTR)

51

2.4.3.2 Fracture theory Several fracture theories are proposed as applicable to the graphite strength. From taking into consideration of simplicity in design and conservative in evaluation, the maximum principal stress failure theory, partially introducing the modified CoulombMohr theory, is adopted on the basis of strength data. As shown in Fig. 2.22 with reference data [23], the maximum principal stress failure theory should be applied in the first quadrant (tensiletensile stress condition). This theory can be even extended partially into the fourth quadrant (tensilecompressive stress condition). In the lower part of the fourth quadrant where the compressive stress component is higher, the modified CoulombMohr fits the data satisfactorily.

2.4.3.3 Stress classification There are two kinds of stress categories, primary stress and secondary stress. The primary stress is produced by pressure load, dead load, etc., and this stress is not reduced by the deformation of components; instead, there is stress redistribution. On the other hand, the secondary stress is produced by thermal load, etc., and this stress is characterized by the stress decrease due to the deformation of components, self-limiting in character. Since the secondary stress is reduced by the plastic deformation for the metallic components, different stress limits for the primary and secondary stresses are regulated, as well known, in the metallic design criteria. On the contrary, graphite, a brittle material with negligible small fracture strain, cannot

Figure 2.22 Biaxial stress theory of PGX graphite [21].

52

High Temperature Gas-cooled Reactors

reduce secondary tress by the plastic deformation. Namely, the secondary stress, for example, thermal stress, will potentially produce cracks and then will produce fracture. The secondary stress is classified in the same manner as the primary stress to evaluate stress limitation conservatively. Namely, stress limitation is taken to the primary plus secondary stress.

2.4.3.4 Stress limit Peak stress is also limited in order to prevent crack initiation and growth even for a single stress (static fatigue). Bending tests show that graphite bars subject to pure bending exhibit strength higher than companion specimen brought to the same maximum stress in uniaxial tension, which can be derived from Weibull theory. A predicted failure by the theory is shown in Fig. 2.23. This suggests that higher stresses may be allowed for tension plus bending stresses than for uniaxial tension stress.

Figure 2.23 Predicted membrane plus bending stress failure curves by the Weibull theory and design stress limits for core support graphite components [21].

Design of High Temperature Engineering Test Reactor (HTTR)

53

Figure 2.24 Design fatigue life diagram of IG-110 graphite [21].

From the facts mentioned above, in the design criteria, the stress limits apply to the three stress components: membrane stress, point stress (membrane plus bending stress), and peak stress (total stress). The design fatigue life diagram is expressed in terms of the ratio of minimum to maximum applied peak stresses on the basis of test data as shown in Fig. 2.24. Several kinds of R ratio show experimental data. For example, 10/3 means 210 MPa compressive stress (minimum stress), and 23 MPa compressive stress (maximum stress). The design life diagram is determined with a statistical basis of 99% survival probability with a 95% confidence level, which is consistent with the basis for the specified minimum ultimate strength. Additionally, in a fatigue evaluation, a linear cumulative damage law is assumed, and the usage factors are limited to 1/3, 2/3, and 1 for operating conditions I and II, III and IV, respectively. Special limits of pure shear stress and buckling stress are also considered in the design criteria. The design stress limits (Hopper diagram) are presented for core support components and core components in Figs. 2.25 and 2.26, respectively.

2.4.3.5 Buckling limit Buckling tests simulating the load imposed on the support post have been conducted to obtain the empirical data for assessing the buckling limit of the support post, because the buckling behavior of the support post is generally difficult to predict only analytical approach. The test results give the formula of critical compressive stress (σcrit) [24], which is the RankineGordon type, as shown in Fig. 2.27. The design critical stress (σd) is determined conservatively from the critical compressive stress, and the design limits are specified in the design criteria. The design limits are determined such that the safety factor in each operational condition is in accordance with that of the design stress limits for membrane primary plus secondary stresses.

54

High Temperature Gas-cooled Reactors

Figure 2.25 Design stress limit for core support graphite components [21].

Figure 2.26 Design stress limit for core graphite components [21].

2.4.3.6 Stress analysis Since the maximum principal stress failure theory, partially introducing the modified CoulombMohr theory, is adopted, the maximum stress is calculated. An elastic analysis is basically carried out in order to estimate the higher stress. The basis for determining stressstrain field is linear elastic stress analysis for the core support graphite components and linear viscoelastic (irradiation-induced creep) stress analysis for the core graphite components.

Design of High Temperature Engineering Test Reactor (HTTR)

55

Figure 2.27 Buckling limit of core support post for each operating condition [21].

Figure 2.28 Viscoelastic stress analysis model [21].

Since the mechanical and physical properties of graphite change with temperature, irradiation, and oxidation, the design criteria specify that the design analysis shall be made in consideration of these changes. The analytical techniques are proposed in the design criteria for evaluation of irradiation and oxidation effects. A viscoelastic mechanical model, as shown in Fig. 2.28, is specified to use in the analysis of the thermal/irradiation environmental effects for the core graphite components [21].

56

High Temperature Gas-cooled Reactors

2.4.3.7 Specified minimum ultimate strength The specified minimum ultimate strength, Su, is determined from statistical treatment of strength data such that the survival probability is 99% at a confidence level of 95% because the ultimate strength exhibits considerable statistical scatter. Table 2.7 shows the Su values for graphites and carbon specified in the design criteria. As the environmental effect on the strength, both strength increase by neutron irradiation and strength decrease by oxidation are considered; however, strength increases by temperature and strain rate are not considered due to a conservative design viewpoint.

2.4.3.8 Oxidation effect The graphite components in the HTTR are subjected to impurity reactants in the helium gas coolant during a normal operation and might react with O2 or H2O in air or water ingress accidents. There are three kinds of reaction regimes, depending on temperature. At low temperature, in the “chemical regime,” the reactions are so low that the reactant can penetrate the graphite in depth, causing rather uniform attack and thus reducing the graphite strength without changing apparent geometries. At high temperature, in the “mass transfer regime,” the chemical reactivity is so high that thinning of component occurs because of successive surface oxidation. Between these two regimes, in the “in-pore diffusion-controlled regime,” the reactants diffuse in the pores of graphite with the gradient of concentration, resulting in the reduction of strength depending on the burnoff profile. Since reaction rates depend on the temperature, the kind of reactants and graphite grade, oxidation analysis must be carried out in detail to estimate the burnoff

Table 2.7 Specified minimum ultimate strength regulated by the graphite design criteria [21].

IG-110

PGX

Level A (MPa)

Level B (MPa)

Level C (MPa)

Tensile strength

19.4

17.6

15.2

Compressive strength

61.3

57.3

51.0

L

6.4

5.9

5.4

T

5.2

4.4

3.4

L

26.6

25.0

23.0

T

26.1

25.0

23.0

L

4.9

4.4



T

4.8

4.2



L

46.9

43.6



T

41.6

39.2



Tensile strength

Compressive strength

ASR0RB

Tensile strength

Compressive strength

L, axial direction for original block; T, radial direction for original block; Level A, survival probability 99% with confidence level of 95%; Levels B and C, applicable to low stress components.

Design of High Temperature Engineering Test Reactor (HTTR)

57

profiles of graphite structures. The strength of oxidized graphite is specified in the design criteria to be evaluated in the following manner: 1. Geometry reduction. The region where amount of oxidation exceeds the 80% burnoff shall be regarded as burned away completely from the structure. 2. Strength reduction. The tensile and compressive strength decreases of grade IG-110 are shown as a function of burnoff [25] in Fig. 2.29A and B, respectively [21]. Similar plots are shown in Fig. 2.30A and B for the PGX graphite [21,26,27]. The stress evaluation shall be made according to these figures.

Figure 2.29 Dependence of strength on burnoff in uniformly oxidized IG-110 graphite [21]. (A) Tensile strength. (B) Compressive strength.

Figure 2.30 Dependence of strength on burnoff in uniformly oxidized PGX graphite [21]. (A) Tensile strength. (B) Compressive strength.

58

High Temperature Gas-cooled Reactors

2.4.4 Quality control Industry-wide standards for raw material formulations and processing of graphite and carbon have not been established. Therefore graphite and carbon must be selected on the basis of the appropriate required by reactor design. Their quality should be controlled during production and after machining with respect to their properties and specifications. Fig. 2.31 shows the flow diagram of acceptance test program with the manufacturing process of graphite’s and carbon materials.

2.5

Metallic components

2.5.1 Introduction The primary cooling system components and related components that serve as the reactor coolant pressure boundaries of the HTTR are used at high temperatures in

Figure 2.31 Flow diagram of acceptance test program for graphite and carbon [21].

Design of High Temperature Engineering Test Reactor (HTTR)

59

creep regime. In particular, the heat transfer tubes and hot header of the IHX are subjected to temperatures above 900 C. The RPV as well as the metallic core support structures are exposed to the reactor coolant at temperatures of around 400 C under an irradiation condition. High temperature structural materials are chosen for the high temperature components of the HTTR, taking into careful considerations the service conditions and safety functions of the components. The materials used are as follows: G

G

G

G

nickel-base corrosion and heat-resistant superalloy, Hastelloy XR; normalized and tempered (NT) 21/4Cr1Mo steel; two types of austenitic stainless steel, SUS321TB and SUS316; and 1Cr0.5MoV steel, an alloy steel bolting material for high temperature service.

Components and their service conditions are listed up for these structural materials in Table 2.8 [13]. Some of the high temperature materials and their service temperatures are beyond the well-established high temperature structural design codes such as the Elevated Temperature Structural Design Guide for the Prototype Fast Breeder Reactor “Monju” (abbreviated as FBR Code) and the ASME Boiler and Pressure Vessel Code Case N-47 [28]. Accordingly, development of a new high temperature structural design guideline was necessary for these materials at their service temperatures. Moreover, at the very high temperatures, where creep deformation is significant, component design based on elastic analysis is not possible. Thus extensive R&D was carried out not only in JAERI but also in national and private research organizations in Japan to establish a reliable high temperature structural design guideline.

2.5.2 Development of Hastelloy XR Taking into account service conditions of the IHX of the HTTR shown in Table 2.8, a nickel-base CrMoFe superalloy Hastelloy X, which has excellent accumulated experiences in jet engines, was selected for the heat transfer tubes and the hot header in the IHX. Hastelloy X is specified as SB-435, -572, -619, -622, and -626 for sheets and plates, bars, welded pipes, seamless pipes and tubes, and welded tubes, respectively, in the ASME Boiler and Pressure Vessel Code Section III Division 1. Since Hastelloy X does not have sufficient compatibility with the primary helium coolant at very high temperatures, Hastelloy XR was developed from Hastelloy X to improve the compatibility. It was found that for Hastelloy X, tightening the contents of some elements even within the specification of the chemical compositions results in remarkable improvements in the compatibility. The following modification items (1) and (2) were done to Hastelloy X to improve the compatibility and further modification items (3) and (4) were given to improve the applicability to the HTTR. 1. Optimizing manganese and silicon contents: formation of stable and adherent oxidation films is essential for the very high temperature components. Such an oxidation film is formed on the base metal through optimizing the Mn and Si contents for Hastelloy X [29]. 2. Lowering aluminum and titanium contents: internal oxidation and intergranular attack are suppressed through lowering the Al and Ti contents [29].

Table 2.8 Material and service conditions of HTTR high temperature components [13]. Material

9/4Cr1Mo steel

Hastelloy XR

Product form

Plate, forging, pipe

Tube, plate forging

Components

Service conditions Design temperature ( C)

Design pressure (MPa)

Reactor pressure vessel

440

4.8

Shells of intermediate heat exchanger, primary pressurized water cooler, etc.

430

4.8

Outside pipe of concentric double pipe

430

4.8

Intermediate heat exchanger heat transfer tubes

955

0.29

Intermediate heat exchanger hot header

940

0.29

Maximum allowable temperature ( C)

550

1000

SUS321

Tube

Primary pressurized water cooler heat transfer tubes

380

4.8

650

SUS316

Bar

Core restraint mechanism

450



650

1Cr0.5MoV steel

Forging

Core restraint mechanism

450



450

Design of High Temperature Engineering Test Reactor (HTTR)

61

3. Lowering cobalt content: radioactive contamination in the primary cooling system by co-containing corrosion products decreases to negligible levels through lowering the Co content [29]. 4. Optimizing boron content: addition of boron improves the creep strengths for Hastelloy XR [30] but causes contamination of the core and degradation in weldability. Optimization of the boron content is needed for a specific purpose. To a Tungsten-arc Inert-gas (TIG) welding wire, addition of boron within 4060 ppm was made to improve the creep strengths of the welded joints.

The specifications of Hastelloy X and the improved version of Hastelloy X, called as the nuclear grade alloy Hastelloy XR, are shown in Table 2.9. Fig. 2.32 shows results of long-term corrosion tests under severe thermal cycles, wherein superiority of Hastelloy XR to Hastelloy X is demonstrated as expected from the protective oxide film formed on Hastelloy XR [31].

2.5.3 Identification of failure modes A high temperature structural design guideline provides design limits and rules for guarding high temperature components against failure modes. Development of a new high temperature structural design guideline, therefore, requires; 1. Identification of failure modes under exposure to service environments within the guideline application temperature range for each material and 2. development of design limits and rules for guarding against each failure mode with appropriate safety margins.

From reviewing material test results and information on failures at commercial plants and experimental facilities, the following failure modes were identified for the five structural materials mentioned above: 1. 2. 3. 4. 5. 6. 7.

ductile rupture by short-term loading, creep rupture by long-term loading, buckling by short-term loading, creep buckling by long-term loading, creepfatigue failure, gross distortion by incremental collapse and ratcheting, and loss of function by excessive deformation.

These failure modes are the same as those considered in well-established hightemperature structural design codes. It should be noted here that the long-term loading means loading at high temperatures that develops significant creep effect over a long period.

2.5.4 Developments of design limits and rules The fact that the failure modes for the new materials are the same as for those of the well-established codes suggests the possibility that fundamental philosophies on design limits and rules of the well-established codes can be applicable to the new materials. Among the well-established high temperature structural design codes, the

Table 2.9 Specifications for chemical compositions of Hastelloy XR and X [31]. Material

Chemical components (wt.%) Range

Elements C

Mn

Si

P

S

Cr

Co

Mo

W

Fe

Ni

B

Al

Ti

Cu

Hastelloy XR

Maximum

0.15

1.00

0.50

0.040

0.030

23.00

2.50

10.00

1.00

20.00

Remainder

0.010

0.05

0.03

0.50

Minimum

0.05

0.75

0.25





20.50



8.00

0.20

17.00

Remainder









Hastelloy X

Maximum

0.15

1.00

1.00

0.040

0.030

23.00

2.50

10.00

1.00

20.00

Remainder

0.010

0.50

0.15

0.50

Minimum

0.05









20.50

0.50

8.00

0.20

17.00

Remainder









Design of High Temperature Engineering Test Reactor (HTTR)

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Figure 2.32 Comparison of environmental effect (Cr-depleted zone depth) between Hastelloy XR and X [31].

FBR Code was the only one that had been authorized by the Japanese government, and so it was the most appropriate for discussion on applicability to new materials. We came to the conclusion that design limits and rules for the abovementioned seven failure modes and for the five materials can be developed on the basis of the fundamental philosophies of the FBR Code. On this conclusion, the detailed design limits and rules were developed for each material, based on experimental data on material properties and structural mechanics behavior under multiaxial stress states, referring to those of the FBR Code, as described below.

2.5.4.1 Hastelloy XR 2.5.4.1.1 Material characterization

The maximum metal temperature of Hastelloy XR in the HTTR reaches about 900 C even during the normal operation and is likely to exceed 950 C but less than 1000 C in events such as a loss of secondary cooling. Taking into account the service temperature conditions, material tests and structural mechanics tests for both base metals and TIG-weld joints were conducted at temperatures ranging from room temperature to 1050 C, mainly in JAEA but also in the National Research Institute for Metals and research laboratories of private nuclear power companies. Test conditions of major material property tests for the base metals are briefly listed in Table 2.10. Test specimens were taken from product forms of tubes, plates, forging cylinders, and bars simulating application to the HTTR high temperature components. By carefully reviewing the experimental data, the following results were derived.

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High Temperature Gas-cooled Reactors

Table 2.10 Mechanical properties data on Hastelloy XR obtained for high temperature structural design [29]. Test item

Test conditions

Tensile tests

Temperatures: room temperature to 1000 C, every 25 C Strain rates: 0.3100 %/min

Creep tests

Temperatures: 500 C1050 C, every 50 C Maximum test time: about 38,000 h Total number of tests: about 300

Fatigue and creepfatigue interaction tests

Temperatures: room temperature to 1000 C, every 50 C at high temperatures Strain rates: 2 3 1025 to 1 3 1023 s21 Hold times: 01 h Materials: as received and thermally aged

Fracture toughness tests

Thermal aging conditions Temperatures: 800 C1000 C Maximum aging time: 2000 h Test items: V-notch charpy, fracture toughness, and fatigue crack propagation rate

Corrosion tests

Environment: HTTR coolant gas-simulated helium Temperatures: 900 C1000 C Maximum test time: 30,000 h

Others

2.5.4.1.2 Tensile property

Poisson’s ratio, thermal expansion, and so on

At low or intermediate temperatures up to 800 C, Hastelloy XR is work hardening under monotonic loadings at the strain rate of 0.3 %/min, which is specified for tensile tests by the Japanese Industrial Standards (JIS), and has hardening ratios of two or above, similarly to austenitic stainless steels (the hardening ratio is defined as a ratio of ultimate tensile strength to yield strength). On the other hand, at high temperatures above 850 C, an abrupt decrease in load or a wavy stressstrain curve under straining at this strain rate is observed due to dynamic recrystallization. Taking into consideration that dynamic recrystallization is not observed at higher strain rates of about 100 %/min as shown in Fig. 2.33, the strain rate for the tensile tests is changed for Hastelloy XR from 0.3 to 100 %/min at high temperatures over

Design of High Temperature Engineering Test Reactor (HTTR)

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Figure 2.33 Stressstrain curve for Hastelloy XR (1000 C, extension rate 5 100 %/min) [31].

800 C. The time-independent allowable limits were generated from the tensile test data at this higher strain rate.

2.5.4.1.3 Creep property Since many commercial superalloys are known to lose their stability of mechanical strength at very high temperatures above 1000 C, the maximum temperature by which high temperature strengths, in particular creep rupture strength, are stable for Hastelloy XR is required to be identified. Fig. 2.34 shows that trends in stress dependence and data scattering of the creep rupture strength are quite similar at 1000 C to those at lower temperatures. Therefore it was concluded that Hastelloy XR is stable up to 1000 C. Concerning the helium environmental effect on creep rupture strength, Fig. 2.35 shows creep rupture lives under a specific stress in various helium environments. Hastelloy XR suffers no degradation in creep rupture strength except in a decarburizing environment. In this figure, a helium environment is characterized fairly well in the stability diagram for Cr (aCr 5 0.8), which is expressed by a carbon activity ac and oxygen partial pressure PO2. Atmospheres denoted as the areas I and II lead to rapid decarburization with or without oxidation, while in the areas IV and V rapid carburization occurs. In area III, mild carburization occurs. In Fig. 2.35, a creep rupture life at a specified helium environment is scaled to lengths of the bar located on the stability diagram. A detailed description of this diagram is given in Kurata et al. [32]. The primary coolant of the HTTR shall be in the area III where any significant degradation in creep rupture life is not observed for Hastelloy XR. Then, it is not necessary to consider helium environment effects on design allowable limits for Hastelloy XR.

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High Temperature Gas-cooled Reactors

Figure 2.34 Creep rupture life for Hastelloy XR [31].

Figure 2.35 Comparison of creep rupture lives for Hastelloy XR in several different helium environments on the stability diagram for Cr (aCr 5 0.8) at 950 C under 26 MPa [31].

2.5.4.1.4 Creepfatigue interaction Creepfatigue interaction for Hastelloy XR is quite similar to those for austenitic stainless steels such as SUS304 and 316. Degradations in lifetime are more pronounced due to holds in tension than those in compression. Fig. 2.36 shows the applicability of the well-known cycle and time fraction rule proposed by Robinson

Design of High Temperature Engineering Test Reactor (HTTR)

67

Figure 2.36 Strain rate effect and hold time effect on creepfatigue interaction for Hastelloy XR [31].

[33] and Taira [34], which is adopted in the FBR Code. It can be concluded form these figures that the linear summation rule of cycle and time fractions is applicable to Hastelloy XR with a great deal of safety margins even at very high temperatures.

2.5.4.1.5 Applicability of the fast breeder reactor code As discussed above, the material properties for Hastelloy XR were observed to be basically similar to those for austenitic stainless steels. These observations lead to the conclusion that the FBR Code is, in principle, applicable to Hastelloy XR at the temperatures ranging to 1000 C, with a modification to the tensile test procedure.

2.5.4.1.6 Structural mechanics behavior The high temperature structural design guideline for class 1 components of the HTTR was established on the basis of component-wise structural mechanics behavior data as well as material property data referring to the FBR Code. The emphasis of the structural mechanics research works was placed on the applicability of the FBR Code to Hastelloy XR under the service conditions of the very hightemperature components. Research works for Hastelloy XR include experiments on multiaxiality of creep rupture strength and creepfatigue interaction, and on creep buckling. Further research works were carried out for establishment of creep analysis methods for Hastelloy XR.

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High Temperature Gas-cooled Reactors

2.5.4.1.7 Multiaxiality of creep rupture strength and creepfatigue damage Since the very high temperature components are exposed to multiaxial loading conditions, multiaxial formulations are required for high temperature strengths of Hastelloy XR. In the FBR Code, the stress intensity criterion, that is, the maximum shear stress criterion, is adopted as the multiaxial formulation for primary stresses, while Von Mises’ stress is that for the primary 1 secondary stresses in evaluating a creep damage. Experiments were carried out in such manners that a tubular test specimen was subjected to a combination of axial and torsional loads. From the experiments, it was concluded that the Von Mises’ criterion predicts the creep rupture life on the safe side. Consequently, the multiaxial formulations, which were given in the FBR Code, were demonstrated to be applicable to Hastelloy XR.

2.5.4.1.8 Creep buckling Heat transfer tubes of the IHX shall not fail by a creep buckling at a piping rupture accident in the secondary cooling system. Component-wise experiments, therefore, were conducted at the HENDEL at JAEA so as to demonstrate the structural integrity of the tubes against the creep buckling and the applicability of a design rule given in the design code. The creep bucking data demonstrated the structural integrity with a great safety margin. A finite element calculation predicted the creep buckling time in good agreement with the experimental data, that is, within an accuracy of 50%.

2.5.4.1.9 Creep analysis method Key items for establishing an appropriate creep analysis method are as follows: 1. generation of an appropriate creep constitutive equation, 2. definition of correct safety margins for uncertainties in predicting creep behavior, and 3. a procedure to define loading sequences or combinations.

For the item (1), several research experiments were carried out to clarify a hardening rule and a flow rule under multiaxial stress states and also statistical analyses were made to formulate a creep equation, that is, a correlation of creep data from constant uniaxial load tests under isothermal conditions. The experimental data showed the applicability of strain hardening rule and Von Mises’ flow rule to Hastelloy XR. The statistical analyses revealed that the time function proposed by Garofalo et al. [35] correlates the creep curve data in the superior agreement to the rational time function [36]. For the item (2), principles to define the safety margins for variations in creep behavior of a high temperature structure were established through sensitivity analysis of a creep constitutive equation [37]. The analytical results clarified that the variations might be covered with those in fundamental creep property such as creep strain curves. For the item (3), creep analyses of the very high temperature components were conducted, taking into account a unique feature of thermal transient behavior of the components. Finally, the design limits and rules for Hastelloy XR in the HTTR high temperature structural design guideline were developed referring to those of the FBR Code, with exceptions.

Design of High Temperature Engineering Test Reactor (HTTR)

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2.5.4.2 21/4 Cr1Mo steel 21/4Cr1Mo steel has a variety of applications to pressure-retaining components in service at elevated temperatures, in particular to boilers and liquid metal fast breeder nuclear reactors (LMFBRs), including the Japanese prototype LMFBR “Monju” For this material, the FBR Code specifies that design rules and limits are applicable to class 1 components. The scope of the FBR Code for 21/4Cr1Mo steel is as follows: G

G

G

maximum application temperature: 550 C, product forms: tube and plate, and specified environment: sodium but no neutron irradiation.

In the HTTR, the NT material is applied to the RPV as well as the piping, shells of heat exchangers, etc. of class 1 components. During normal operation, the RPV attains a temperature of about 400 C; it is exposed to the highest temperature of about 530 C maximum during severe accidents such as a rupture of the primary concentric hot gas duct. The RPV is also exposed to neutron irradiation even during normal operation but the accumulated neutron flux is less than 1 3 1017 n/cm2 (E . 1 MeV), which is negligible. Product forms of this material are pipes, plates, and forgings in application to the HTTR class 1 components. Among these products, the plates and forgings are out of scope of the well-established Japanese structural design standards, entitled “Technical standards for LWR power plant components—MITI standards No. 501.” The service conditions of high temperature and neutron irradiation require tightened specifications for chemical composition contents of impurities such as silicon and sulfur for guarding against brittle fractures due to thermal neutron irradiation embrittlement. This requirement is met by imposing additional requirements on the chemical composition specifications of the JIS. Many mechanical strength tests as well as fracture toughness tests have been done to clarify the following fundamental characteristics for 21/4Cr1Mo steel of the HTTR specification: 1. 2. 3. 4.

mechanical strengths as compared to those of materials of the untightened JIS specifications, mechanical strengths among the product forms of tubes, plates, pipes, and forgings, thermal aging and/or neutron irradiation effects on mechanical strengths, and simulated HTTR-helium gas environment effects on mechanical strengths.

Discussion on these mechanical strength data came to the following conclusions: 1. Mechanical strengths for the HTTR specification material are equivalent to those of the JIS specification material for creep rupture strength. 2. Among the product forms, the following relationships regarding mechanical strength exist: tube 5 pipe plate 5 forging. 3. Thermal aging and/or neutron irradiation do not cause any significant degradation in mechanical strengths in the service ranges of high temperatures up to 550 C and of a neutron flux up to 1 3 1019 n/cm2. 4. Simulated HTTR-helium also does not cause any degradation below a high temperature of 550 C.

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High Temperature Gas-cooled Reactors

Accordingly, design rules and limits for the HTTR specification material were generated in the HTTR high temperature structural design guideline from those for the JIS specification material, which are specified in the FBR Code.

2.5.4.3 Austenitic stainless steels SUS321TB and SUS316 SUS321TB and SUS316 are both well-experienced heat-resistant materials with superior mechanical strengths at temperatures up to about 700 C and are included in the FBR Code. In the HTTR, SUS321TB is applied to heat transfer tubes of helium-to-PWCs, that is, of the PPWC, the SPWC, and the auxiliary heat exchanger. The tube maximum temperature is about 400 C during normal operation. SUS316 is applied to reactor internal structures such as a core restraint mechanism, and control rod guide tubes. The maximum temperatures of these structures are also about 400 C during normal operation. Service conditions of the abovementioned structures are within the scope of the FBR Code with only one exception of the HTTR-helium gas environment effect. Such an effect is negligible at normal operating temperatures of the structures. Accordingly, design rules and limits for the both stainless steels were generated in the HTTR high temperature structural design guideline from those specified in the FBR Code.

2.5.4.4 1Cr0.5MoV steel 1Cr0.5MoV steel has widespread application for bolts for chemical plant vessels and is used as a bolting material in the MITI standard No. 501 for nuclear application below a maximum temperature of 425 C. For its application to high temperature components, design rules and limits have been specified up to 600 C in the JIS B 8243 “Construction of Pressure Vessels” but are not included in well-established nuclear high temperature structural design codes such as the FBR Code. In the HTTR, 1Cr0.5MoV steel (JIS SNB16) is applied to not only bolts for vessels such as the RPV and gas circulator castings but also for parts of the core restraint mechanism. A service temperature of 1Cr0.5MoV steel is around 400 C, a little higher than its creep regime threshold temperature of about 380 C under which creep and relaxation effects are negligible. This means that 1Cr0.5MoV steel is served under a “semicreep” condition. Many mechanical strength tests were carried out to qualify the semicreep temperature condition. It has been clearly found from these tests that creep effects are not significant up to 450 C for creep rupture strength but are significant for higher temperature around 500 C for fatigue strength. Then, it was defined that the semicreep temperature range for 1Cr0.5MoV steel should be 375 C450 C. Taking into consideration such limited service temperatures of semicreep condition in the HTTR, structural design rules and limits for 1Cr0.5MoV steel are developed through essentially extrapolating design rules and limits for application under low temperature condition, where creep deformation and damage are negligible.

Design of High Temperature Engineering Test Reactor (HTTR)

2.6

71

Core components and reactor internals

2.6.1 Introduction The reactor consists of core components, reactor internals, reactivity control equipment, and the RPV. One column is the row of prismatic hexagonal blocks piled up axially. The active core consists of 30 fuel columns and 7 control rod (CR) guide columns, and is surrounded by 12 replaceable reflector columns, 9 reflector region CR guide columns, and 3 irradiation test columns, as shown in Fig. 2.4 [13,38,39]. The permanent reflector blocks, surrounding replaceable reflector blocks, are fixed by the core restraint mechanism. The reactor internals consist of graphite and metallic core support structures and shielding blocks as shown in Fig. 2.7. They support and arrange the core components, such as fuel elements and replaceable reflector blocks within the RPV. The graphite core support structures consist of permanent reflector blocks, hot plenum blocks, support posts, CBSs, etc. The metallic core support structures consist of support plates, a support grid, a core restraint mechanism, etc. This chapter describes the core components and reactor internals design of the HTTR and the program of in-service inspection (ISI) to confirm structural integrity of the reactor internals.

2.6.2 Fuel The fuel element of the HTTR is a so-called pin-in-block-type fuel element, which is composed of fuel rods and a hexagonal graphite block. The fuel rod is classified into four types in the HTTR. One is A-type fuel rod, which is used as a driver fuel. The others are B-type fuel rods, namely, B-1, B-2, and B-3, which have different specification of coating layers of CFPs and are used in irradiation tests for advanced fuels. The configuration of the fuel element is shown in Fig. 2.6. The CFP consists of a microsphere of low-enriched UO2 with the TRISO coating. The CFPs are incorporated into fuel compacts with a graphite matrix. The fuel rod, which is composed of fuel compacts and a graphite sleeve, is contained within a vertical hole of a graphite block.

2.6.2.1 Design requirement In the fuel safety design of HTGR, it is important to retain FPs within the CFPs so that their release to the primary coolant may not exceed an acceptance level. From this point of view, the basic design criteria for the fuel are to minimize the failure fraction of as-fabricated fuel coating layers and to avoid significant additional fuel failures during operation. To meet the latter criterion, the fuel temperature is limited below 1495 C during normal operation conditions and below 1600 C during anticipated operational occurrences, and the fuel burnup is limited to 33,000 MWd/t based on the results of irradiation tests [40]. On these basic considerations, the safety requirements for the HTTR fuel, except for the graphite block, were settled as follows [41]: 1. The initial failure fraction of the as-fabricated coating layers of the CFPs shall be less than 0.2% in terms of the sum of heavy contamination and SiC defects, while the

72

2.

3.

4.

5.

High Temperature Gas-cooled Reactors

expected fraction is less than 5 3 1024. The value of 0.2% was determined from the viewpoint to limit off-site exposure during normal operation. The CFPs shall not fail systematically under normal operating conditions. a. The penetration depth of PdSiC interaction shall not exceed the thickness of the SiC layer of 25 μm, because the fully penetrating PdSiC interaction will lead to a loss of the FP retention function of the SiC coating layer. b. The kernel migration shall not exceed the thickness of 90 μm, which is the sum of the first and second layers. The CFPs shall be designed so as to avoid failure considering irradiation-induced damage and chemical attack through the full-service period; namely, the additional failure fraction in the coating layers of the CFPs shall be less than 0.2% through the full-service period. The value of 0.2% was determined in the same manner as that for the initial failure fraction. The maximum fuel temperature shall not exceed 1600 C at any anticipated operational occurrence to avoid fuel failure. The temperature criterion is established to avoid any significant failure and remarkable degradation in the coating layers of the CFPs, taking into account the frequency of the anticipated operational occurrences and those continuous times at high temperature. The behavior of the irradiated CFPs was examined in a temperature range up to 2400 C, with a furnace installed in a hot cell in order to ensure fuel integrity at elevated temperature [42]. The temperature criterion of 1600 C was determined on the basis of these test results [13]. In addition to above requirements, the following requirements were settled to guarantee mechanical integrity of the fuel compact and the graphite sleeve. a. The fuel compact and the graphite sleeve shall not be broken or cracked considering thermal stress and irradiation-induced damage. b. The fuel compact and the graphite sleeve shall not contact with each other to keep their mechanical integrity.

2.6.2.2 Design details The specification of the CFPs is shown in Table 2.11. Table 2.12 shows the specification of the fuel compact, the graphite sleeve, and the fuel block. The CFP consists of spherical fuel kernel of low-enriched UO2 (12 kinds of enriched UO2 from 3.4% to 9.9% and about 6% on the average) with the TRISO coating. The TRISO coatings consist of a low-density, porous PyC buffer layer (60 μm) adjacent to the fuel kernel (600 μm in diameter) followed by high-density isotropic pyrolytic carbon layer, an SiC layer (25 μm) and a final outer PyC coating. The CFPs are incorporated with graphite matrix into the fuel compact, which is 10 mm in inner diameter, 26 mm in outer diameter, and 39 mm in height. The fuel rod, which is composed of fuel compacts and the graphite sleeve, is contained within a vertical hole of a graphite block. Helium gas flows downward through the 3.5 mm annular gap between the vertical hole and the fuel rod to remove heat produced by fission and gamma heating.

2.6.2.3 Evaluation The fabrication of the first-core fuel started in June 1995 and took in total 33 months. A total of 66,780 fuel compacts in 126 fabrication lots were produced corresponding to 4770 fuel rods, using a total of 900 kg of uranium. The fabrication data show the low

Design of High Temperature Engineering Test Reactor (HTTR)

73

Table 2.11 Specification of coated fuel particles [38]. Item

A-type

B-1/B-2 type

B-3 type

Fuel type

Rod

Rod

Rod

Fuel coating type

TRISO

TRISO

TRISO

Diameter of particle (μm)

920

940

830

Material

UO2

UO2

(U, Th)O2(Th/ U 5 4)

Density (% of T.D.)

95

95

95

Diameter (μm)

600

570

500

First layer

Low-density PyC 60

Low-density PyC 80

Low-density PyC 60

Second layer

High-density PyC 30

High-density PyC 30

High-density PyC 30

Third layer

SiC 25

(SiC (B-1))/(ZrC (B-2)) 35

SiC 30

Fourth layer

High-density PyC 45

High-density PyC 40

High-density PyC 45

310 (average 6)

5

20

Fuel kernel

Material and thickness (μm)

Enrichment of 235U (wt.%)

initial failure fraction of the order of 1025 as compared with the requirement of 0.2% [43]. For the PdSiC interaction in the HTTR fuel, a relationship between the maximum penetration depth and the amount of Pd released from the kernel was expected by cubic root [44], from which the maximum penetration depth of the HTTR fuel through the full-service period is estimated. About 11 μm of the penetration depth for the maximum is obtained, which is far less than the safety design limit of the total SiC layer thickness of 25 μm [45]. The distance of the kernel migration in the HTTR fuel was calculated based on the R&D results. The maximum kernel migration length is 55 μm, which is far less than the safety design limit of 90 μm. From the tests on fuel behavior under accidental conditions, it was found that the coating layers of the HTTR-CFPs would maintain their intactness below 1600 C within the range of HTTR design condition, such as burnup [13,40]. The requirements of mechanical integrity of the fuel rods were verified by irradiation tests, in which neither crack nor break of the fuel rods was observed in the irradiation range. It is also verified through extensive R&D that the B-type fuel elements will meet the design requirements and fulfill their function at normal operation as well as anticipated operational occurrences.

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High Temperature Gas-cooled Reactors

Table 2.12 Specification of fuel compact, graphite sleeve, and fuel block [38]. Item

Properties

Fuel compact Type

Hollow cylinder

Material

Coated fuel particle, binder, and graphite

Packing fraction of coated fuel particle

30 vol.% (A and B-3); 35vol.% (B-1 and B-2)

Dimension Outer/inner diameter (mm)

26/10

Height of a compact (mm)

39

Effective length of a fuel rod (mm)

546 (14 fuel compacts)

Graphite sleeve Type

Cylinder

Material

Graphite

Dimensions Outer diameter (mm)

34

Thickness (mm)

3.75

Length (mm)

580

Gap width between fuel compact

0.25 and graphite sleeve (mm)

Fuel block Type

Pin-in-block

Configuration

Hexagonal

Dimensions Width across the flats (mm)

360

Height (mm)

580

Fuel hole diameter (mm)

41

Material

Graphite

Number of fuel rods in a block

33 or 31

2.6.3 Hexagonal graphite blocks The core is an array of hexagonal graphite blocks made up of fuel blocks, control rod guide blocks and replaceable reflector blocks. These blocks provide the structure for the arrangement and confinement of the fuel material; neutron moderation, heat

Design of High Temperature Engineering Test Reactor (HTTR)

75

transfer, and positioning of control/shielding absorber materials. The core consists of vertical columns of hexagonal blocks arranged on a uniform triangular pitch. Within the array of the vertical columns, 30 columns contain fuel. The vertical structure of the column is composed of nine hexagonal graphite blocks. The triangular pitch of the columns on each support block is 362 mm at cold condition. The hexagonal fuel block is 360 mm in width across the flats and 580 mm in height. The core is approximately 2.3 m in equivalent diameter and 2.9 m in height.

2.6.3.1 Design requirement The design requirements for the hexagonal graphite blocks are as follows: 1. The fuel and BP rods are retained within the core at all the operational conditions. 2. Coolant channels remain free of obstructions and offset displacement, which would impair the core cooling capability. 3. Control rod insertion holes remain free of obstructions and offset displacements, which would impair the insertion of control rods. 4. Reserved shutdown system pellet insertion holes remain free of obstructions and offset displacement, which would disturb the insertion of B4C/C pellets or would cause a loss of absorber pellets from the core. 5. The blocks support the structure located above. 6. The capability to handle blocks using the fuel handling machine is not impaired. 7. The integrity of the fuel and BP rods or shield pins is not impaired.

A loss of structural integrity is defined as any damage to the blocks or dowel/ socket connections, which would prevent any of the above requirements from being met. The structural integrity of the graphite blocks is assured by limiting the maximum stress to the value given in the Graphite Structural Design Guideline of core components [10]. The calculated stresses should include the effects of dead weight, pressure and seismic loads, and thermal and irradiation-induced strains. Creep and changes in physical properties as a function of temperature and fast neutron fluency should be considered as well as the effect of chemical reactions with oxidizing impurities contained in the coolant.

2.6.3.2 Design details The primary objective of the mechanical core design is to provide a structurally stable core array, which satisfies the nuclear, thermal/hydraulics, seismic design, and refueling operation requirements. The location and structural restraint for the columns in the core are provided by the core support structures. The core columns have structural features to maintain the alignment of the coolant channels and control rod insertion holes, to ensure a proper coolant flow distribution, and also to ensure that the control absorbers can be inserted as required. All graphite blocks are designed to support the weight of blocks above, loads created by coolant flow pressure differences, seismic loads, loads induced by core restraint, and fuel handling forces. Internal stresses due to dimensional changes induced by thermal and irradiation gradients are considered in the block structural design.

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High Temperature Gas-cooled Reactors

The fuel blocks are graphite hexagonal prisms, which are 360 mm in width across the flats and 580 mm in height, with arrays of coolant channels and BP insertion holes as shown in Fig. 2.6. Thirty-one or thirty-three fuel channels of 41 mm in diameter extend through the block and are aligned with coolant channels in the blocks above and below. The fuel rod is supported in each channel and is cooled by helium flowing through the annular gap. Three threaded dowels are installed on the top face coupled with sockets in the bottom face of the block located above. A hole at the center of each fuel block is provided for fuel handling. The hole profile is shaped so that a lifting ledge is machined at the lower end. Additional holes are provided in the corners of the blocks for the insertion of the BP rods. Control rod guide blocks with control rod and reserved shutdown system pellet insertion holes are of the same external shape and envelope dimensions as the fuel blocks and use the same dowel/socket connections as shown in Fig. 2.37. These blocks have two control rod insertion and one reserved shutdown system pellet insertion holes of 123 mm in diameter. Envelope dimensions and dowel pattern of the block are the same as the fuel block. The bottom block of the control rod guide column contains B4C/C pins for the thermal neutron shield. The top and bottom replaceable reflector blocks have the same basic configuration as the fuel blocks but do not contain fuel rods. These reflector blocks, above and below the active core, have the same arrangement of coolant channels as the fuel blocks within the same columns. The side replaceable reflector blocks adjacent to the core have the same envelope dimensions as fuel blocks but are solid graphite and contain only a central handling hole for removal and insertion. The bottom replaceable reflector block below each fuel block column provides a transition of the many coolant channels to a single large channel, which mates with the coolant Control rod insertion hole Reserved shutdown pellet insertion hole

Dowel pin

Graphite block

Dowel socket

Figure 2.37 Schematic view of control rod guide block [38].

Design of High Temperature Engineering Test Reactor (HTTR)

77

channels within the hot plenum blocks. The alignment of the fuel, control rod guide, and replaceable reflector columns with the hot plenum blocks is maintained by dowel/socket connections in the bottom reflector blocks, which fit into the hot plenum blocks. The bottom layer of reflector blocks contains B4C/C pins for the thermal neutron shield.

2.6.3.3 Evaluation The structural integrity of the graphite components is ascertained by limiting the calculated maximum principal stress to the value defined in the graphite structural design criteria [22]. The characteristics of the design criteria are presented in this issue [46]. The basis for determining stressstrain fields is linear viscoelastic (irradiation-induced creep) stress analysis for core graphite components. The stress analysis, thus, is carried out by a specially developed viscoelastic stress analysis code VIENUS [47]. The fuel blocks are categorized into three column regions, taking into account of the similarity of the thermal and boundary conditions, as shown in Fig. 2.38. The critical stress condition is the combined residual stress, caused by both irradiation-induced creep and dimensional change, and transient thermal stress. With increasing the neutron irradiation, tensile residual stress increases at the central region of the fuel blocks, while compressive at the periphery region. When inner coolant channels are suddenly cooled down in the transient condition, tensile thermal stress arises at the center region and compressive thermal stress periphery region. This transient event is a severe condition for the fuel blocks, and the lifetime of the fuel block is limited from it. The maximum residual stresses and transient stresses, occurred in the fuel blocks as shown in Fig. 2.38, are shown in previous study [48]. The maximum transient stress appears at the third-layer C column block, while the maximum residual stress appears at the fourth-layer C-column block. This is due to the different temperature boundary conditions for each fuel block position; the cold gas flows downward at the top part with increasing its temperature. The operational stress and residual stress at third-layer C-column are plotted in Fig. 2.39. The operational stresses are reduced by the creep deformation; on the other hand, the residual stresses are increased due to the accumulation of residual strain by the irradiation-induced creep deformation and dimensional

A column B column C column

Figure 2.38 Arrangement of fuel blocks in HTTR.

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High Temperature Gas-cooled Reactors

Figure 2.39 Operational stress and residual stress at the third-layer C column [38].

Figure 2.40 History of calculated maximum stress in the fuel block [38].

change. The combined stresses are plotted in Fig. 2.40. The transient stresses increase at the first stage until about 2400 s from the beginning of the transient, and then these stresses reduce gradually. The maximum point plus peak stress component is about 12 MPa, which is about a half of the limiting stress. The stress analysis for the control rod guide blocks is performed using the VIENUS code under the same condition as the fuel blocks. Table 2.13 shows estimated stress based on the graphite structural design criteria for the control rod guide block. The maximum tensile stress of the membrane component is about 81% of the stress limit as seen in Table 2.13. The stress analysis for the replaceable reflector blocks is also performed under the same conditions as for the fuel blocks. The bottom replaceable reflector block just below the active core gives the highest stress among

Design of High Temperature Engineering Test Reactor (HTTR)

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Table 2.13 Stress evaluation of control rod guide block [38]. Operation condition

Stress componenta

Stress

b

Membrane

Point

Peak

6.0, 25.3

6.2, 25.8

6.4, 26.4

I and II

Estimated

III

Limited

7.4, 220.3

11.1, 230.4

20.1, 254.7

Estimated

6.0, 25.3

6.2, 25.8

6.4, 26.4

Limited

11.3, 230.7

16.7, 245.6

20.1, 254.7

Estimated

6.0, 25.3

6.2, 25.8

6.4, 26.4

Limited

15.8, 243.0

20.1, 254.7

22.3, 260.7

IV

a

Positive and negative values mean tensile and compressive stresses, respectively. Stresses are calculated conservatively during over cooling, which leads to the maximum estimated stress components in plant condition I. b

the replaceable reflector blocks. The maximum tensile stress of the membrane stress component is about 78% of the stress limit for the bottom replaceable reflector block.

2.6.4 Core support structures The core support graphite structures mainly consist of permanent reflector blocks, hot plenum blocks, support posts, and bottom structures as shown in Fig. 2.37. The permanent reflector blocks surrounding the replaceable reflectors are made up of large polygonal graphite blocks fixed by key elements and a core restraint mechanism. The support posts located between the hot plenum blocks and CBSs provide a hot plenum space where the hot core outlet helium gas can be mixed uniformly. To design these graphite core support structures, the graphite structural design code was developed on the basis of graphite test data. During an earthquake, key/keyway structures and support posts are the important components to maintain the core array. In order to obtain design data such as design load to verify the seismic response and to develop the seismic analysis code, seismic tests, therefore, were performed using two kinds of scaled models, a 1/5-scale model simulating a horizontal two-dimensional array of the hot plenum blocks and a 1/3-model simulating a horizontal two-dimensional array of CBSs including the support post structures [49].

2.6.4.1 Design requirement The design requirements for the graphite core support structures are summarized as follows: 1. To have sufficient strength to maintain the structural integrity during normal operation and anticipated operational occurrences, and to maintain the structural functions during an earthquake or accidents.

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High Temperature Gas-cooled Reactors

2. To ensure the core cooling capacity by the restriction of leakage flow through the permanent reflector blocks and hot plenum blocks. 3. To support the core array during an earthquake and to control the mechanical interaction of the core components so that these will not fail. 4. To ensure the coolability of the metallic core support structures.

2.6.4.2 Design details The hexagonal hot plenum block array is made up of two axial layers. The upper layer consists of sealed plenum blocks, in which leakage flow between these blocks is sealed by the triangular seal elements, and the lower layer is composed of keyed plenum blocks, which provide lateral and vertical positioning and support of the core array. The hot plenum block assembly contains passages for the primary coolant flow from the columns in the core to the hot plenum. These blocks are fabricated from grade PGX graphite, a medium-to-fine grained, molded structural graphite. Severe load is applied to the hot plenum block during an earthquake. The design seismic load is determined by the vibration test of core support structures as mentioned in the following section. The stress analysis for the hot plenum block is carried out by the finite element stress analysis code. The estimated stresses based on the graphite structural design code are shown in Table 2.14. The maximum tensile peak stress is about 79% of the limited peak stress at the S2 earthquake, an extreme design earthquake. The support posts and seats are so designed as to support the core and hot plenum block array while providing the hot plenum for the primary coolant flow. The support posts and seats are made of grade IG-110 graphite. The maximum stress for the support post and seat component arises at the spherical contact area. The evaluated stresses at this point during an earthquake are summarized in Table 2.15. The maximum compressive peak stress component becomes about 83% of the limited peak stress during the S1 earthquake, the maximum design earthquake.

Table 2.14 Estimated earthquake stresses based on the graphite structural design code for the hot plenum block [39]. Operation condition

Stress

Stress componenta (MPa) Membrane

I, II, and III and S1

I, II, and III and S1

a

Point

Peak

Estimated

1.0

2 0.9

1.8

2 1.9

2.6

2 2.1

Limited

2.2

2 8.8

2.9

2 11.8

4.0

2 15.9

Estimated

1.2

2 1.3

2.3

2 2.7

3.5

2 2.9

Limited

2.7

2 10.6

3.5

2 14.1

4.4

2 17.6

Positive and negative values mean tensile and compressive stresses, respectively.

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Table 2.15 Estimated earthquake stresses based on the graphite structural design code for the support post [39]. Operation condition

Stress

Stress componenta (MPa) Membrane

I, II, and III and S1

I, II, and III and S1

Estimated

1.6

2 18.8

4.6

Limited

9.7

2 30.7

Estimated

1.6 11.6

Limited a

Point

Peak 234

12.4

2 45.8

12.9

2 40.9

17.5

2 55.2

2 19.4

4.6

2 34.2

12.4

2 45.8

2 36.8

15.5

2 49.1

19.4

2 61.4

Positive and negative values mean tensile and compressive stresses, respectively.

Table 2.16 Estimated earthquake stresses based on the graphite structural design code for the permanent reflector [39]. Operation condition

Stress

Stress componenta (MPa) Membrane

Point

Peak

I, II, and III and S1

Estimated

0.7

22.2

1.4

23.4

2.5

25.5

111

Limited

1.7

28.8

2.3

211.8

3.1

215.9

I, II, and III and S1

Estimated

0.9

22.1

1.9

23.5

3.1

25.6

111

Limited

2.3

210.6

2.7

214.1

3.4

217.6

a

Positive and negative values mean tensile and compressive stresses, respectively.

The CBSs consist of three blocks: lower plenum blocks, carbon blocks, and bottom blocks. These blocks have a function of thermal insulation between the hot plenum and metallic core support structures. The permanent reflector is a graphite structure immediately surrounding the replaceable reflector and control rod guide columns. It is an assembly of graphite blocks consisting of 12 circumferential segments in 8 axial layers fabricated from grade PGX graphite. The severe stress of the permanent reflector is created by the earthquake load at the connecting keyway position. The evaluated stresses based on the graphite structural design code are summarized in Table 2.16. The maximum tensile peak stress reaches about 90% of the limited value for the S2 earthquake. The typical properties of graphite and carbon are listed in Table 2.17. These components are so designed as to maintain the structural integrity during accidents and earthquakes in accordance with the graphite structural design code considering the brittle nature of graphite material.

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Table 2.17 Typical material characteristic values for graphite and carbon (unirradiated material) [39]. Item

IG-110 graphite

PGX graphite

ASR-0RB carbon

Bulk density (mg/m3)a

1.78

1.73

1.65

25.3

8.1

6.8

76.8

30.6

50.4

7.9

6.5

8.7

Mean thermal expansion coefficient (293673 K) (1026 K)

4.06

2.34

4.40

Thermal conductivity (W/(m K)) (673 K)

80

75

10

Ash (ppm)

Max. 100

Max. 7000

Max. 5000

Grain size (μm)

Mean: 20

Max. 800

Max. 2000

Mean tensile strength (MPa)a Mean compressive strength (MPa)

a

Young’s modulus (GPa) (( 6 1/3)Su) a

b

a

At room temperature. Determined from the cord joining two points (one point is the one-third of the specified minimum tensile strength and the other is the one-third of the specified minimum compressive strength) on the stressstrain curve. b

2.6.4.3 In-service inspection and surveillance test In order to confirm structural integrity of the core support graphite structures, visual inspection will be performed using a TV camera as ISI and measurement of material characteristics using Surveillance Test specimen (Surveillance Test). Fuel replacing and ISI will be carried out at the same period, and the surveillance test will be carried out at the second and fourth refueling time and at the end of the plant lifetime.

2.6.4.3.1 In-service inspection using TV camera To conduct the visual inspection, the ISI system was developed. Fig. 2.41 shows the structure of the ISI system. In the visual inspection, surface flaws as well as an array of core support graphite structures will be visually examined. The required flaw size in the visual inspection is determined on the basis of the fracture mechanics approach taking into consideration of the fracture toughness and stress profiles in the graphite structures. Possible visual inspection areas by a TV camera are inner surfaces of the permanent reflector blocks, upper surfaces of the sealed plenum blocks, side surfaces of core support posts, and upper surfaces of the lower plenum blocks. The important positions of core support graphite structures in the visual inspection are as follows: G

Permanent reflector blocks surface flaws on the permanent reflector blocks gaps between the permanent reflector blocks G

G

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Motor

Wire roller drum

Cable roller drum Cable Wire Support unit Guide roller

TV camera

Figure 2.41 Schematic view of ISI system [39]. G

Hot plenum blocks gaps between the permanent reflector blocks and hot plenum blocks surface flaws on the hot plenum blocks Hot plenum region inclination of the support posts surface flaws on the lower plenum blocks gaps between the lower plenum blocks G

G

G

G

G

G

2.6.4.3.2 Results of preservice inspection In order to obtain the reference data of ISI, a preservice inspection (PSI) of core support graphite structures was carried out using the developed ISI system. Fig. 2.42 shows one of the results of the PSI, the surface of permanent reflector blocks, and the axial gap between permanent reflector blocks. No harmful flaw and abnormal array were observed in the PSI. In the visual inspection, structural integrity of the core support graphite structures will be evaluated by comparing the results of PSI.

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High Temperature Gas-cooled Reactors

2.6.4.3.3 Surveillance test Measurement of material properties using surveillance test specimens will be carried out to evaluate the irradiation effect on mechanical and thermal properties of graphite and carbon. Table 2.18 shows the measured properties in the surveillance test. Bending and compressive strengths, Young’s modulus, etc. will be measured at regular intervals. An oxidation weight loss will be also measured to examine the oxidation-induced damage.

Figure 2.42 Surface of permanent reflector blocks and axial gap between permanent reflector blocks [39].

Table 2.18 Measured properties in the surveillance test [39].

Permanent reflector block

Test items

Material

Dimension change

PGX

Bending strength Hot plenum block

Bending strength

PGX

Oxidation weight loss Support post

Young’s modulus Bending strength Compressive strength Oxidation weight loss

IG-110

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2.6.5 Core support metallic structures The core support metallic structures are composed of the support plate, the core support grid, the core restraint mechanism, etc. as shown in Fig. 2.7. The support plate and core support grid are located under the core bottom insulation layer. The core restraint mechanism surrounds the permanent reflector blocks.

2.6.5.1 Design requirement The support plate and core support grid is so designed as to form a statically stable foundation for the core components and reactor internals under the highest postulated temperatures and under seismic loads. The core restraint mechanism has to fulfill the following functions: 1. To provide a stable lateral support not only during normal operation condition in the anticipated operational occurrences but also during accident conditions and the seismic conditions for the full core height. 2. To prevent radial outward movement of the permanent reflector blocks in order to avoid forming excessive bypass gaps between blocks. 3. To reduce impact loads on core components, reactor internals during a seismic event. 4. To maintain sufficient annular space for the primary coolant flow between the permanent reflector blocks and the RPV.

2.6.5.2 Design details The support plate forms a plain foundation surface for the core components and reactor internals, and is composed of steel plates with a thickness of 89 mm. The support plates are set on steel support posts, guiding the core weight to the core support grid below. The core support grid transfers the total weight of the core to the RPV through the support ribs welded on the inner surface of the hemispherical bottom head closure. The grid is reinforced with welded diamond shaped plate structures. The support plate and the core support grid are cooled by the primary coolant flowing into the RPV. The core restraint mechanism surrounds and stresses the permanent reflector blocks with 10 axially distributed units. Each unit consists of 12 restraint bands, 12 band supports, 2 restraint rings, and radial keys as shown in Fig. 2.43. The restraint bands are stressed to produce radial force on the band supports that is transmitted to the permanent reflector blocks via the side shield blocks and support legs. The restraint rings, located above and below the restraint bands, are normally stress-free. They limit the radial displacement of the bands during a seismic event and restraint the lateral load of the core directly in case the restraint bands fail. In such a case, the restraint rings transfer the load to the RPV through the radial keys. Normally, there is no contact between the restraint rings and the radial keys. The radial keys are individually adjusted before being attached to the RPV and are provided to compensate for circumferential fabrication tolerances.

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High Temperature Gas-cooled Reactors

Permanent reflector block Reactor pressure vessel Support leg Coolant path Radial key Radial key sheet Band support Restraint band Restraint ring Side shielding block

Figure 2.43 Arrangement of core restraint mechanism [39].

Compression element

Tube

Tensile element

Rod

Figure 2.44 Schematic view of core restraint band [39].

The restraint bands are composed of a central rod and six concentric pipes as shown in Fig. 2.44 except for the 9th and 10th units. The 9th and 10th units have eight concentric pipes. The pipes are connected in series with the pipes mutually acting as compression and tensile elements. With the bolt eyes connected to the central bar on one end and to the outer pipe on the other end of the band, the whole arrangement forms a sort of tensile spring. The component materials of the restraint band are SNB16 steel for the tensile element and SUS316 for the compression element. Even though the thermal expansion coefficients of both steels are different from those of graphite and core structural materials, the coefficient of the band itself is adjusted to be equal to that of

Design of High Temperature Engineering Test Reactor (HTTR)

87

graphite. Therefore the restraint of the permanent reflector blocks is maintained in spite of change in temperature. The restraint mechanism is used during plant lifetime without repair. Relaxation of SNB16, used for the tensile element, under high temperature creep conditions causes reduction of the restraint force that could even lead to coolant leakages through gaps between the permanent reflector blocks. For the bottom layer of the band, at the end of plant lifetime, approximately 60% of restraint force is reduced by the relaxation of SNB16 with a conservative estimation. This result satisfies design requirement of the restraint band. Therefore mechanical properties of SNB16 are important and were accumulated for the design of the restraint mechanism.

2.6.5.3 In-service inspection and surveillance test The restraint mechanism is used during plant lifetime without repair. By the relaxation of SNB16, used for the tensile element, under high temperature creep conditions, the restraint force will decrease, and then coolant leakages through gaps between the permanent reflector blocks will increase. Therefore deformation characteristics of SNB16 are important, and the relaxation as well as Charpy impact tests of the SNB16 are programmed during HTTR operation as surveillance tests. These tests will both be carried out after 5 and 10 years of operation and at the end of the plant lifetime.

2.6.6 Shielding blocks Shielding blocks consist of top and side shielding blocks and are composed of a neutron-absorbing material and casing. The top shielding block is installed at the top of each column and has coolant channels matching the channels in the column below. The side shielding blocks are installed outside of the permanent reflector blocks and compressed inward by the core restraint mechanism.

2.6.6.1 Design requirement Requirements for the design of the shielding blocks are as follows: 1. The shielding blocks have the function to limit the thermal neutron fluence in the metallic core support structures and the RPV. 2. The top shielding blocks for the fuel column have coolant channels of adequate diameter to adjust coolant flow rate distribution. 3. The side shielding blocks shall have sufficient strength to transfer compression force by the core restraint mechanism to the permanent reflector blocks.

2.6.6.2 Design details The top shielding block is a hexagonal block, which is composed of a neutronabsorbing material of sintered B4C/C, a casing of SUS316 and dowel pins. The top shielding blocks for the fuel and control rod guide columns have coolant channels matching the channels in the succeeding blocks. The top shielding block for the fuel

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High Temperature Gas-cooled Reactors

column has an internal cavity for coolant in order to prevent neutron streaming. Fig. 2.45 shows a schematic structure of the top shielding block for the fuel column. Dowel pins on the top are used for positioning of the handling head of fuel handling machine. The side shielding block is also composed of sintered B4C/C and a SUS316 casing as shown in Fig. 2.46. The compressive force by the restraint band is transferred to the side shielding blocks through the band supports as shown in Fig. 2.43. Four support legs are connected directly to the thick SUS316 plate, which locates RPV side of the block, and transfer the compressive forces to the permanent reflector blocks. Handling hole

Dowel pin

Coolant channel Casing

Neutron absorber (B4C/C)

Figure 2.45 Schematic view of top shielding block [39].

Casing

Neutron absorber (B4C/C)

Support leg

Figure 2.46 Schematic view of side shielding block [39].

Design of High Temperature Engineering Test Reactor (HTTR)

2.7

89

Seismic design

2.7.1 Introduction The seismic design for the HTTR facilities was based on “Guidelines for Aseismic Design of Nuclear Power Plants (GAD)” [50]. It is stated in the guidelines that important buildings and structures must be supported on base rock. The reactor building of the HTTR, however, has been planned to be supported on a sand layer formed during the Quaternary era. For this reason, various seismic safety evaluations for the HTTR facility were investigated for construction of the HTTR building [51]. After the HTTR was constructed, seismometers were installed in the surrounding foundation and in the reactor building for confirmation of the vibration behavior of the HTTR. The seismic response analysis was performed using a seismic observation record and basic earthquake motion [51]. Fig. 2.47 shows a vertical view of the core, the CBS, and RPV. Permanent disarray of, or damage to, the core components induced by an earthquake would prevent insertion of control rods. Aseismic design of the core, thus, is one of the major concerns in the safety evaluation of the HTTR. In the earthquake, the HTTR core requires

Figure 2.47 Vertical view of reactor internals and vessel [52].

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High Temperature Gas-cooled Reactors

structural integrity to maintain the safety functions of reactor shutdown and decay heat removal. Since the core is composed of piled graphite blocks with their associated clearances on the CBS, the vibration response of the blocks is highly nonlinear with multiple impacts occurring between adjacent blocks [52]. Aseismic studies of the core were carried out through experimental and analytical approach to clear the singularity of vibrational behavior due to impact phenomena [48,53]. The basic guidelines of seismic design, geological composition, R&D results on aseismic studies, and the typical integrity evaluation result of graphite components of the HTTR [3,4] are described.

2.7.2 Seismic design 2.7.2.1 Basic guideline of seismic design The nuclear power plant is so designed that all anticipated earthquakes may not cause a severe accident. Then the structures are considered to possess seismic safety on the basis of following provision. 1. The core, which is a multilayered structure of graphite blocks, is supported on the metallic core support structure. Buildings, structures, systems, and equipment except for the core should be made rigid as a rule. Moreover, important buildings and structures should be supported on a foundation, which is safe enough against any postulated earthquake. 2. The degree of importance of facilities is classified into classes A, B, and C from viewpoint of effect of radiation release to the environment. The seismic design should be done according to this degree of importance. 3. Abovementioned facilities of classes A, B, and C should be designed so as to resist against a seismic force, which is obtained from the story shear coefficient according to the degree of importance. 4. Facilities of class A should be so designed that they may resist seismic forces determined by use of a dynamic analysis with the basic earthquake motion S1 defined according to GAD. 5. Especially, important facilities among those of class A are classified into “class As.” They must maintain a safety function basic earthquake motion S2 defined according to GAD. Moreover, a dynamics analysis is done for equipment and piping of class B, which might vibrate in resonance. 6. Seismic forces to facilities of the class A act on a disadvantageous direction, which is a combination of a horizontal force obtained from basic earthquake motion with a vertical force obtained from a vertical seismic coefficient which is half of the maximum amplitude acceleration of seismic earthquake motion. It is assumed that the vertical earthquake coefficient is constant in the direction of height. 7. If there is a possibility that components and piping resonate with the structure, dynamic analysis should be carried out.

2.7.2.2 Seismic classification The classification of facilities is shown in Table 2.19. 1. Class A: Facilities storing radioactive material, or relating directly to the facilities which stores radioactive material and may release radioactive material to the environment by the loss of function, and/or one to prevent these situations and one to mitigate the effect of released radioactive material.

Table 2.19 Classification of HTTR facilities [51]. Class

Facilities belonging to each class

Main facilities

Seismic classification

A

(1) Equipment and pipings composing a reactor coolant pressure boundary

Reactor pressure vessel

As

Vessels, pipings, gas circulators, and valves belonging to reactor coolant pressure boundary

As

(2) Facilities to store spent fuel

Spent fuel storage pool

As

Spent fuel storage rack

As

(3) Facilities to add negative reactivity for the emergency shutdown of a reactor and to maintain the subcritical condition

Control rod and control rod drive mechanism (as for scramability)

As

(4) Facilities to remove decay heat from a core after the emergency of a reactor)

Auxiliary cooling system (except an inner piping of concentric hot gas duct

As

(5) Facilities to remove decay heat from a core after a failure of reactor coolant pressure boundary

Reactor vessel cooling system

As

(6) Facilities act as pressure barrier in case of failure of a reactor coolant pressure boundary and to prevent directly the release of radioactive material

Reactor containment vessel

As

Pipings and valves belonging to the reactor containment vessel boundary

As

(7) Facilities to control the release of radioactive material to environment in case of the accident with such possibility except facilities belonging to “6”

Emergency air purification system

A

(8) The other

Core internal structure (except facilities of class As)

A

Reserved shutdown system

A

Primary helium purification system (except facilities of classes B and C)

A

Fuel failure detection system (except facilities of classes B and C)

A

Primary helium sampling system (except facilities of classes B and C)

A

(Continued)

Table 2.19 (Continued) Class

Facilities belonging to each class

Main facilities

Seismic classification

B

(1) Equipment connecting directly to reactor coolant pressure boundary which are storing primary coolant or capable to store it

Primary helium purification system (except facilities of classes A and C)

B

Fuel failure detection system (except facilities of classes A and C)

B

Primary helium sampling system (except facilities of classes A and C)

B

(2) Facilities storing radioactive waste except facilities, which have a little storage or less effect of radiation to public by the failure than the limitation of dose equivalent at outside of surrounding observed area

Radioactive waste disposal system (except facilities of class C)

B

(3) Facilities related to radioactive material except waste and ones which may give excessive radiation exposure by a failure to the public and the worker

Fuel handling machine

B

Ceiling crane

B

Shield with much reduction of radiation

B

Reactor pressure vessel leak detection system

B

Ceiling crane in the spent-fuel storage building

B

(4) Facilities to cool spent fuel

Pool water purification system for spent pool (a part related to cooling water pool)

B

(5) Facilities to control a release of radioactive material in case of an accident and not belonging class A

A part of ventilation and air conditioning system in the spent fuel storage building

B

(Continued)

Table 2.19 (Continued) Class

Facilities belonging to each class

Main facilities

Seismic classification

C

(1) Equipment not belonging to classes A and B

Control rod drive mechanism (except a part related scramability)

C

Secondary helium sampling system

C

Fresh fuel storage system

C

Secondary helium cooling system

C

Pressurized water-cooling system

C

Primary helium purification system, primary helium sampling system, fuel failure detection system not related to high radioactive material

C

Secondary helium purification system

C

Primary helium storage and supply system

C

Secondary helium storage and supply system

C

Waste disposal system not related to high radioactive material

C

Pool water-cooling and purification system (related to the supply of pool water)

C

Fire extinguishing system

C

Ventilation and air conditioning system

C

Air compressioning system for nonsafety use

C

Cooling water system for nonsafety use

C

The others

C

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High Temperature Gas-cooled Reactors

2. Class B: Above-mentioned facilities, the facilities of which the effect and the influence are comparatively small. 3. Class C: Facilities for which the safety equal to the conventional industrial facilities is required.

2.7.2.3 Basic design earthquake ground motion Basic design earthquake ground motions are defined at the free field of the base stratum surface in the HTTR site. The base stratum is a firm base, which was formed in the tertiary or earlier geological era and, which is not weathered significantly with shear wave velocity of more than 700 m/s. At the HTTR sites, the level is set up to be the upper surface of the Neogene-Tertiary Miocene based on the survey results. The basic earthquake motions were defined, as shown in Fig. 2.48, by means of Ohsaki spectra. At this time, the maximum design earthquake S1 and the extreme design earthquake S2 are supposed based on the investigation on the historical earthquakes, the active faults, and the seismotectonic structure. The maximum accelerations of S1 and S2, which are envelop curves of these earthquakes, are 1.8 m/s2 (180 gal) and 3.5 m/s2 (350 gal), respectively.

2.7.3 Geological composition and seismometry 2.7.3.1 Geological composition To grasp the geological composition at the HTTR site and obtain samples for laboratory tests, a boring survey was performed at the positions as shown in Fig. 2.49. The geological sections obtained from the survey are shown in Fig. 2.50. The supporting foundation at the HTTR site is formed by horizontal layers with continuity.

2.7.3.2 Seismometry A seismometry system was installed in the HTTR facility to confirm the behavior of a seismic event, and earthquake observation was started in October 1997. Seismometers were installed in the surrounding foundation at the positions of underground 250, 94, 30, and 1 m, and in the reactor building at the position of BF3, BF1 1F and 2F. Figs. 2.51 and 2.52 show the points of seismometers in the surrounding foundation and in the reactor building, respectively.

2.7.4 Structure of core components The reactor core is an array of graphite blocks composed of fuel elements, a control rod guide, and replaceable reflector, which provide neutron moderation and heat transfer, and serves for positioning of control and shielding absorber materials. One column is the row of prismatic hexagonal blocks piled up axially. The active core consists of 30 fuel columns and 7 control rod guide columns and is surrounded by 12 replaceable reflector columns, 9 reflector region control rod guide columns, and

Design of High Temperature Engineering Test Reactor (HTTR)

95

Figure 2.48 Flow diagram of definition of basic earthquake ground motion [51].

3 irradiation test columns. Each column consists of nine piled graphite blocks and one top shielding block. The top shielding block is a hexagonal block, which is composed a neutron-absorbing material of sintered B4C/C, a casing of SUS316. The permanent reflector blocks are fixed by the core restraint mechanism. The graphite block of the fuel elements is a hexagonal right prism with an array of fuel element holes, as shown in Fig. 2.23. The block is 0.36 m across the flats

96

Figure 2.49 Position of geological survey [51].

High Temperature Gas-cooled Reactors

Design of High Temperature Engineering Test Reactor (HTTR)

Figure 2.50 Geological section [51]. (A) EastWest. (B) NorthSouth.

97

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High Temperature Gas-cooled Reactors

Figure 2.51 Position and depth of seismometers in surrounding foundation [51].

Figure 2.52 Position of seismometers in reactor building [51].

and 0.58 m high. Three dowel pins are installed on the top face, coupled with sockets in the bottom face of the block above. The dowel-socket system ensures the correct orientation of the blocks within the column with respect to each other. The control rod guide block shown in Fig. 2.38 and the replaceable reflector block have the same external dimensions as the fuel element block.

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2.7.5 Development of evaluation method Aseismic studies have been performed to clarify the vibration characteristics and assess the structural integrity for the core components and the CBS. These studies have been done independently through experimental and analytical methods.

2.7.5.1 Vibration characteristics of core components A core seismic computer code is required for determination of the overall core response and the displacement characteristics of the core and design loads for the core components. A comprehensive R&D program was conducted to develop a means to analyze the block-type core for a seismic design. However, no generalpurpose codes met the above requirements. This was due to a number of limitations, especially the inability to efficiently treat impacts and highly nonlinear characteristics in large numbers. We, therefore, developed the SONATINA-2V code [54]. This code is used to predict the impact phenomena between graphite blocks and to provide information on impact forces, displacements, etc. that is required for the safety evaluation of the structural integrity of the core. To evaluate the validity of the SONATINA-2V code and to confirm the structural integrity of the core graphite blocks, many seismic tests were conducted.

2.7.5.2 Validation of SONATINA-2V code The impact of the graphite block, particularly the dowel and socket system, and the displacements of blocks were the most important items from the viewpoint of the core seismic design. Therefore the code was verified in terms of these response values [53]. Fig. 2.53 shows the typical overall core response characteristics at 2.5 m/s2. The analytical results were in good agreement with the test. Since the SONATINA-2V code applied to a two-dimensional analytical model, it is necessary to confirm that the code can predict the seismic response of the three-dimensional full core. Fig. 2.54 presents the relative displacements of the core graphite blocks as a function of excitation frequency at 1.0 m/s2, comparing the test results using the full-scale seven-column model with the analytical one. In a lowfrequency range, where the first vibration mode of the column is dominant, the columns vibrate together. When the frequency exceeds the resonance frequency, the columns exhibit different response modes where the relative displacement response is small. As the code models a two-dimensional vertical slice core (three degrees of freedom), the equivalent column stiffness becomes harder than that of a threedimensional core (six degrees of freedom). As a result, the calculated frequency response shifts to a slightly higher frequency region in comparison with the experimental result. It can be, thus, seen that the analytical results showed good agreement with the test.

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High Temperature Gas-cooled Reactors

Figure 2.53 Typical core overall response characteristics at resonance [52].

2.7.6 Structural integrity of graphite components The dowel socket and keykeyway systems are designed such as to transfer the shear force between connected blocks during earthquake [55]. The core support posts are subjected to vertical loads induced by the vertical and rocking motions of the hot plenum block [56]. Consequently, the excessive load concentrates on these connecting elements such as dowel sockets, keykeyway, and support post seat. The assessment of structural integrity for the connecting elements is one of the major concerns in the aseismic design because their fracture may lead to the serious damage to the reactor core. Thus the evaluation results for their structural integrity will be described in the following subsection.

2.7.6.1 Core components The design load acting on three dowel-socket systems in a fuel block was 7.4 kN in the postulated maximum design earthquake denoted by S2 using the SONATINA-2V

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Figure 2.54 Comparison between test and analytical results of relative block displacement (full-scale seven-column test) [52]. (A) Secondary block from top. (B) Fourth block from top.

code. Based on the fracture tests for a dowel-socket system, the safety load factor (fracture load divided by the design load) becomes about 4.6 for the final failure [55]. It can be seen that the dowel-socket system withstands more than three times as much as the seismic load in an S2 earthquake. After the seismic tests using the fullscale seven-column model, including excitation over the S2 level, the graphite blocks were visually inspected, and there had never been any damage in these components.

2.7.6.2 Core bottom structure The CBS is designed to withstand earthquakes such as S1 and S2. The maximum stress on the keyway corner was estimated to be 1.7 MPa from the measured strain under S2. This stress is sufficiently lower than the fracture stress (about 10 MPa) obtained from the full-scale fracture test of the keykeying system. The seismic loads on the core support posts were evaluated from the 1/3-scale model test. The

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High Temperature Gas-cooled Reactors

design load using the test results gets to be 130 kN and results in an estimate of 7 MPa for the axial compressive membrane stress in the core support post [56]. This stress is sufficiently lower than the strength of the core support post (about 55 MPa). After B60 excitation in the 1/3-scale model test, including excitation over the S2 level, the hot plenum blocks and the core support posts were visually inspected. There was no damage in these graphite components.

2.8

Cooling system

2.8.1 Introduction The reactor cooling system consists of the primary cooling system, the secondary helium cooling system, the pressurized water-cooling system, the auxiliary cooling system, and the VCS as schematically shown in Fig. 2.55. The main cooling system consists of the primary cooling, the secondary helium cooling, and the pressurized water-cooling systems [57]. The primary cooling system has two heat exchangers of the IHX and the PPWC. The primary helium gas is transported from the reactor core to the IHX and to the PPWC through the primary concentric hot gas duct. The secondary helium cooling system, mainly consisting of the SPWC, removes the heat from the primary helium gas through the IHX. The pressurized water-cooling system consists of an air cooler and water pumps. The air cooler cools the pressurized water in both the PPWC and the SPWC and transfers the heat from the reactor core to the atmosphere as final heat sink. The auxiliary cooling system is in stand-by during normal operation and starts up to remove the residual heat after reactor shutdown. The VCS runs at rated flow rate during normal operation to cool the biological concrete shield. It cools the

Figure 2.55 Schematic diagram of cooling system for HTTR [57].

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reactor core at accident such as a primary-pipe rupture when the auxiliary cooling system is no longer able to cool the core effectively. The HTTR has two operation modes. One is the single-loaded operation using only the PPWC, and the other is the parallel-loaded operation using both the PPWC and the IHX to cool the primary helium gas from the reactor. In the single-loaded operation, the PPWC removes the heat of 30 MW. In the parallel-loaded operation, the PPWC and the IHX remove the heat of 20 and 10 MW, respectively. The pressure boundary of the main components such as the IHX, the PPWC, the SPWC, the auxiliary heat exchanger, and the primary concentric hot gas duct are served at temperatures slightly above 400 C. The pressure boundary between the primary and secondary helium gases, such as the heat transfer tubes of the IHX, reaches approximately 900 C during the high temperature test operation. To assure the structural integrity of these high temperature components, the following rules are applied: 1. To use several kinds of heat-resistant metallic materials, taking into account the service conditions of components. 2. To establish a reliable high temperature structural design guideline. 3. To apply countermeasures for overcoming the severity of the service conditions.

As for item (1), two commercial materials and a superalloy, which was developed for the HTTR, are used as follows: a. 21/4Cr1Mo steel for the reactor coolant pressure boundary in-service at temperatures of approximately 400 C; b. austenitic stainless steels (SUS321TB and SUS316) for the heat transfer tubes of the PPWC, the SPWC, and the auxiliary heat exchanger; c. Ni-base corrosion and heat-resistant superalloy Hastelloy XR [58] for the heat transfer tubes of the IHX and other reactor coolant pressure boundaries reaching very high temperature of approximately 900 C.

2.8.2 Primary cooling system 2.8.2.1 Primary pressurized water cooler The PPWC is a vertical U-tube type heat exchanger. Fig. 2.56 shows a schematic view of the PPWC and Table 2.20 shows its major specifications. The hot primary helium gas from the inlet nozzle flows horizontally between the baffle plates and cools the outside surface of the heat transfer tubes, flowing upward and turns backward several times. It flows out via the upper or lower outlet nozzles to the primary gas circulators and flows back to the annular space between the inner and outer shells to cool them. The pressurized water of 3.5 MPa is led to each heat transfer tube and heated up by the primary helium gas. The thermal insulation is installed inside the inner shell so as to maintain its temperature lower than 440 C. A tube sheet supports the heat transfer tubes. The heat capacity of the PPWC can be changed from 30 to 20 MW and vice versa by changing helium gas flow paths according to the loop operational modes. The

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High Temperature Gas-cooled Reactors

Figure 2.56 Schematic view of primary pressurized water cooler [57].

primary helium gas flows out through three lower outlet nozzles during the parallel-loaded operation, while three upper nozzles during the single-loaded operation. The inner and outer shells are made of 21/4Cr1Mo steel. The austenitic stainless steel (SUS321TB), which is superior not only in strength at high temperatures but also anticorrosion, is used for the heat transfer tubes. Material of the baffle plates is Hastelloy XR because they are exposed to the primary helium gas of approximately 950 C for a long time. Burnout of the pressurized water was the major problem to be solved because the maximum temperature difference between the primary helium gas and pressurized water is approximately 800 C, which is higher than that of existing PWCs. To prevent burnout, the flow velocity of the primary helium gas in high temperature region is maintained lower than the average flow velocity in the PPWC. The flow velocity in the low temperature region is maintained higher than the average to promote heat transfer. The size of the PPWC remains compact by the above design method.

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Table 2.20 Major specifications of primary pressurized water cooler [57]. Type

Vertical U-bent tube type

Design pressure Outer shell

4.7 MPa

Heat transfer tube

4.7 MPa

Design temperature Outer shell

430 C

Heat transfer tube

380 C

Operating condition

Rated operation

High temperature test operation

Single-loaded operation

45.2 t/h

37.0 t/h

Parallel-loaded operation

29.7 t/h

24.3 t/h

Flow rate of primary helium gas (max.)



Inlet temperature of primary helium gas

890 C

950 C

Outlet temperature of primary helium gas

395 C

395 C

Single-loaded operation

625 t/h

618 t/h

Parallel-loaded operation

413 t/h

410 t/h

Flow rate of pressurized water



Inlet temperature of pressurized water

135 C

134 C

Outlet temperature of pressurized water

175 C

174 C

Heat capacity Single-loaded operation

30 MW

Parallel-loaded operation

20 MW

Heat transfer tube Number

136

Outer diameter

25.4 mm

Thickness

2.6 mm

Length

10 m

Outer diameter of shell

2.1 m

Total height

7.5 m

Material Outer and inner shell

SCMV4-2NT (21/4Cr1Mo steel)

Heat transfer tube

SUS321TB

Tube sheet

SFVA F22B (21/4Cr1Mo steel)

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High Temperature Gas-cooled Reactors

2.8.2.2 Intermediate heat exchanger The IHX is a vertical helically coiled counter flow-type heat exchanger in which primary helium gas flows on the shell side and secondary helium gas in the tube side as shown in Fig. 2.57. Table 2.21 shows the major specifications of the IHX.

Figure 2.57 Schematic view of intermediate heat exchanger [57].

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Table 2.21 Major specifications of intermediate heat exchanger [57]. Type

Vertical helically coiled counter flow

Design pressure Outer shell

4.7 MPa

Heat transfer tube

0.29 MPa (differential pressure)

Design temperature Outer shell

430 C

Heat transfer tube

955 C

Operating condition

Rated operation

High temperature test operation

Flow rate of primary helium gas (max.)

14.9 t/h

12.2 t/h

Inlet temperature of primary helium gas

850 C

950 C

Outlet temperature of primary helium gas

395 C

395 C

Flow rate of secondary helium gas

12.8 t/h

10.8 t/h

Inlet temperature of secondary helium gas

244 C

237 C

Outlet temperature of secondary helium gas

175 C

174 C

Heat capacity

10 MW

Heat transfer tube Number

96

Outer diameter

31.8 mm

Thickness

3.5 mm

Length

30 m

Outer diameter of shell

1.9 m

Total height

10 m

Material Outer and inner shell

SCMV4-2NT (21/4Cr1Mo steel)

Heat transfer tube

Hastelloy XR

Hot header and center pipe

Hastelloy XR

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High Temperature Gas-cooled Reactors

The primary helium gas enters into the IHX through the inner pipe of the primary concentric hot gas duct. It is deflected under a hot header and discharged around the heat transfer tubes to transfer the heat to the secondary helium cooling system. It flows to the primary gas circulator via the upper outlet nozzle and flows back to the annular space between the inner and outer shells. The secondary helium gas flows downward in the heat transfer tubes and upward in the central hot gas duct through the hot header. The inner insulation is installed inside the inner shell to maintain its temperature below 440 C. The insulations outside and inside the central hot gas duct restrain the heat transfer so that high efficiency can be obtained. In addition, it also keeps the temperature of the central duct below 940 C. The primary helium gas is contained only in the primary cooling system because the pressure in the secondary helium cooling system is adjusted somewhat higher than that in the primary cooling system. The tube support assemblies hold the heat transfer tubes. Both the central hot gas duct and the heat transfer tube support assemblies are hung from the vessel top so that the thermal expansion is not constrained. The material of the heat transfer tubes and the hot header is Hastelloy XR, and the inner and outer shells are made of 21/4Cr1Mo steel. The IHX has a bypass line, which prevents natural circulation from the reactor core to the IHX during the single-loaded operation. The forced circulation from the PPWC through the bypass line occurs and keeps the temperature of the outer shell below 430 C. The primary helium gas flows from the PPWC and enters into annulus space between the inner and outer shells and then flows inside the inner shell through the bypass line. It returns to the PPWC through the IHX and the primary concentric hot gas duct. The shutoff valve stops this forced circulation during the parallel-loaded operation and in case of a scram when the auxiliary cooling system is activated. The inner structures such as the heat transfer tubes, the central hot gas duct, and the hot header are operated beyond 900 C. A design method based on the elastic analysis cannot meet the criteria of the high temperature structural design code for the HTGR class 1 components. Therefore the design method based on a creep analysis is used for evaluation of their structural integrity. The creepfatigue damage was properly evaluated and is capable of meeting the criteria.

2.8.2.3 Primary gas circulator The primary gas circulator is a centrifugal, dynamic gas bearing type circulator. Table 2.22 shows the major specifications of the circulator. Three circulators for the PPWC and one for the IHX are installed in the primary cooling system. The former circulators are operated during both the parallel-loaded operation and singleloaded operation, and the latter during the parallel-loaded operation. During operation, the rotating assembly is fully floating on sets of a dynamic gas bearing system. Fig. 2.58 shows a schematic view of the circulator. The circulator consists of the following components: (1) electric stator and rotor assembly, (2) internal structure supports, (3) thrust bearings and journal bearings, (4) impeller unit, and (5) filter unit. These internal structures are contained in a casing which is cooled by a water jacket. The casing prevents primary helium gas from leaking into

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Table 2.22 Major specifications of primary gas circulator [57]. For the IHX

For the PPWC

Type

Centrifugal gas bearing

Centrifugal gas bearing

Flow rate (max.)

14.9 t/h

15.1 t/h

Head (max.)

79.4 kPa

107.9 kPa

Design pressure

4.7 MPa

4.7 MPa

Design temperature

430 C

430 C

Casing material

SCMV4-2NT SFVA 22B (21/4Cr1Mo steel)

SCMV4-2NT SFVA 22B (21/4Cr1Mo steel)

Type

Cage-type induction motor

Cage-type induction motor

Power

260 kW

260 kW

Revolution speed

300012,000 rpm

300012,000 rpm

Type of frequency converter

Thyristor convertor

Thyristor convertor

Type

Sintering metal

Sintering metal

Material

SUS316

SUS316

Motor

Filter

the atmosphere. The flow rate of primary helium gas is controlled by a variable speed motor using a frequency converter. The filter unit, which is on the top of the circulator, protects the impeller and rotating shaft from dust. The material of the casing is 21/4Cr1Mo steel.

2.8.2.4 Primary concentric hot gas duct The primary concentric hot gas duct consists of an outer pipe, an inner pipe, and a thermal insulator as shown in Fig. 2.59. Table 2.23 shows the major specifications. The cold helium gas of 400 C flows in an annular path, inside the inner pipe the temperature is 950 C. The outer pipe can contain high-pressure helium gas of 4.0 MPa. The inner pipe, which separates the high and low temperature helium gas paths, supports the pressure difference between the high and low temperature helium gas. The pressure difference is about 0.1 MPa. The liner forms a high temperature helium gas boundary and reinforces the ceramic fiber insulator. The internal insulator between the liner and inner pipe minimizes heat loss from the high to low temperature helium gases and maintains the temperature of the inner pipe below 440 C.

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High Temperature Gas-cooled Reactors

Figure 2.58 Schematic view of primary gas circulator [57].

The outer pipe and the inner pipe are made of 21/4Cr1Mo steel, the material of the liner is Hastelloy XR, and the insulation material is a ceramic fiber composed of SiO2 and Al2O3.

2.8.3 Secondary helium cooling system The secondary helium cooling system consists of a secondary gas circulator, the SPWC, and a secondary helium piping. This system is operated during the parallelloaded operation. The heat is transferred from the primary to the secondary helium gas through the IHX. This transported heat is sent to the pressurized water-cooling system through the SPWC. In the future, the hydrogen production system will be connected to the secondary helium cooling system [59].

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Figure 2.59 Cross-sectional view of primary concentric hot gas duct [57].

Table 2.23 Major specifications of primary concentric hot gas duct [57]. Design pressure Outer pipe

4.7 MPa

Design temperature Outer pipe

430 C

Dimension of outer pipe Outer diameter

863.6 mm

Thickness

42 mm

Dimension of inner pipe Outer diameter

660.4 mm

Thickness

15 mm

Thickness of thermal insulator

90 mm

Material Outer pipe

SCMV4-2NT SFVA 22B (21/4Cr1Mo steel)

Inner pipe

SCMV4-2NT (21/4Cr1Mo steel)

Liner

Hastelloy XR

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High Temperature Gas-cooled Reactors

2.8.3.1 Secondary pressurized water cooler The structure of the SPWC is fundamentally the same as that of the PPWC. The SPWC is a vertical U-bent type heat exchanger and has double shells. Inside of the inner shell, it is thermally insulated. 21/4Cr1Mo steel of SCMV4-2NT is used for the inner and outer shells, SUS321TB is used for heat transfer tubes, and Hastelloy XR is used for the baffle plates and the liner.

2.8.3.2 Secondary gas circulator The structure of the secondary gas circulator is nearly the same as that of the primary gas circulator. The secondary gas circulator has one gas circulator, which is vertical centrifugal type. The circulator is composed of the casing, an impeller, a rotor, bearings, and a motor. The speed of the circulator is changed by a frequency converter to control the helium flow rate.

2.8.3.3 Secondary helium piping The secondary helium piping consists of a concentric hot gas duct connecting the IHX and the SPWC, and a single wall piping connecting the SPWC and the secondary gas circulator. Thermal insulator is attached on the inside of the inner pipe of the concentric hot gas duct. The hot helium gas from the IHX flows inside the inner pipe. The helium gas, pressurized by the circulator, flows in the annular space between the inner and outer pipes. The 21/4Cr1Mo steel is used in the inner and outer pipes, and Hastelloy XR is used as the liner.

2.8.4 Pressurized water-cooling system The pressurized water-cooling system is installed to cool primary and secondary helium gas in the PPWC and the SPWC, respectively. The heat is finally released to the atmosphere via the air cooler. This system consists of a pressurized water pump, an air cooler, and piping. The pressure of water is controlled to be lower than that of the primary cooling system in order to minimize the amount of water ingress into the primary cooling system in case of a heat transfer tube rupture accident but still high enough to prevent boiling.

2.8.4.1 Pressurized water pump Two pressurized water pumps (including a spare) of horizontal centrifugal type are installed. They have a capacity of 640 t/h flow rate at 90 m delivery head.

2.8.4.2 Air cooler The air cooler consists of finned heat transfer tubes and blowers. It has a 30 MW cooling capacity. Its air flow rate is 2600 t/h.

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2.8.5 Residual heat removal system The main cooling system is also used to remove the decay heat of the core at normal reactor shutdown conditions. Besides the main cooling system, the HTTR has two other residual heat removal systems, which are the auxiliary cooling system and the VCS. The auxiliary cooling system removes the residual heat during anticipated operational occurrences and for accidents as reactivity insertion or pipe rupture in the secondary cooling system. The VCS removes the residual heat during the loss of coolant accident in the primary cooling system.

2.8.5.1 Auxiliary cooling system The auxiliary cooling system consists mainly of an auxiliary heat exchanger, an auxiliary gas circulator, and an auxiliary water-cooling system including an air cooler. The auxiliary cooling system has a capacity of about 3.5 MW. The system starts up automatically when the reactor is scrammed in case of an accident if the core cooling by a forced circulation is still possible, while the main cooling system is stopped. The auxiliary cooling system has engineered safety features with redundant dynamic components such as gas circulators, water pumps, and valves, which are also backed up with an emergency power supply. The residual heat of the core can be removed even by the VCS without the auxiliary cooling system. The auxiliary cooling system, however, is needed from the viewpoint of operation flexibility because it takes a very long time for cooling down the core only with the VCS.

2.8.5.2 Vessel cooling system The VCS consists of upper, lower, and side cooling panels around the RPV and cooling water circulation systems. The amount of the heat removal is less than 0.6 MW to accomplish the outlet gas temperature of 950 C because the heat removal of the VCS means heat loss from the RPV. Besides, the VCS is required to remove more than 0.3 MW since the temperatures of the fuel and the RPV did not exceed their limit temperature of 1600 C and 550 C, respectively, to keep their structural integrity. The system is used as a residual heat removal system when the forced circulation in the primary cooling system cannot be maintained due to the rupture of the inner pipe or both pipes in the concentric hot gas duct. The VCS has engineered safety features, that is, it is equipped with two independent sets, which are backed up with an emergency power supply. It is operated even in normal operation in order to cool the biological concrete wall shield.

2.9

Reactivity control system

2.9.1 Introduction The reactivity control system of the HTTR is comprised of a control rod system and a reserve shutdown system. Reactivity is controlled by 16 pairs of control rods, which are individually moved by control rod drive mechanisms located in stand pipes connected to

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High Temperature Gas-cooled Reactors

Irradiation hole Control rod guide block Reserved shutdown pellets insertion hole Control rod insertion hole

Fuel region Replaceable reflector region Neutron detector insertion hole

Replaceable reflector block Permanent reflector block Fuel element Irradiation test column Side shielding block

Reactor pressure vessel Core restraint mechanism

Figure 2.60 Horizontal arrangement of HTTR core [60].

the hemispherical top head closure of the RPV. The control rods are inserted into appropriate holes in the core and the replaceable reflector regions as shown in Fig. 2.60 [60]. The control rod drive mechanism inserts and withdraws a pair of control rods using an AC motor. In the event of a scram, which needs to separate the clutch gear teeth, the control rods are inserted into the core by gravity. Ferritic superalloy Alloy 800H is selected for a material of metallic parts of the control rod. The maximum allowable temperature for the control rod to be used repeatedly after scrams is 900 C, that is, the control rod must be replaced when the temperature exceeds 900 C [13]. Since the maximum temperature of the graphite blocks in the fuel region reaches about 1100 C at “high temperature test operation of 950 C reactor outlet gas temperature,” a two-step control rod insertion method for reactor scram must be adopted for the HTTR. The outer nine pairs of control rods in the reflector region are inserted into the core immediately at a scram, and the other inner seven pairs in the fuel region are inserted 40 min later or when the outlet coolant temperature becomes less than 750 C. Preliminary temperature analysis revealed that in most events that cause scrams, the maximum temperature of the control rods is lower than 900 C [61]. The reserve shutdown system is located in stand pipes accompanied by the control rod system. In case the control rods cannot be inserted, the reserve shutdown system drops B4C/C pellets into the core to shut down the reactor.

2.9.2 Control rod system 2.9.2.1 Design requirement The following criteria are established for design of the control rod system: 1. The control rod system shall be designed to shut down the reactor reliably and safely when required during normal operation conditions and accidents including earthquakes as well as at the lowest shutdown temperature of 27 C (300 K). 2. The control rod system shall be capable of controlling the reactivity changes due to temperature and Xenon density, fuel burnup, experimental samples, etc. 3. The control rod system shall be designed to be failsafe.

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4. The reactivity worth and reactivity insertion rate by a control rod shall be restrained to the level that break of reactor internals preventing the core cooling should not occur in a postulated reactivity insertion accident. Even if a stand pipe break should occur, control rod ejection shall be limited by the stand pipe fixing device.

2.9.2.2 Design details The specifications of the control rod system are shown in Table 2.24. Table 2.24 Specifications of HTTR control rod [60]. Items

Properties

Control rod Type

Double circular cylinders with lid and vent

Number

16 pairs (32 rods) (15 pairs in irradiation tests using the center column)

Total length (m)

3.1

Outside diameter (mm)

113

Inside diameter (mm)

65

Sleeve Thickness (mm)

3.5

Material

Alloy 800H

Neutron absorber Outside diameter (mm)

105

Inside diameter (mm)

75

Material

Sintered compact of B4C/C

Spine Diameter (mm)

10

Material

Alloy 800H

Control rod drive mechanism Drive method Normal operations

Rolled up and down by AC motor through control rod support cable

Scram

Separate electromagnetic clutch and the control rod fall by gravity

Number

16 (15 in irradiation tests using center column)

Drive speed (mm/s)

From about 110 (variable)

Insertion time (s)

,12 (after being triggered)

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High Temperature Gas-cooled Reactors

2.9.2.2.1 Control rod The control rod system of the HTTR is shown in Fig. 2.61. To reduce thermal stresses in the sleeves, the neutron absorber sections and spine are separated from each other and each section is sustained by a support ring. For the same purpose, the absorber section is assembled from parts with gaps between them and all parts are mechanically connected without welds. Each absorber section contains a guide ring to restrain them from rolling too much. The lower end of the spine is attached to a tubular-type shock absorber. At the bottom of the core are graphite dishes, also called shock absorber, which absorb impact of a control rod, when the control rod support cable breaks. During normal operation, the temperature of the clad material is about 600 C maximum, whereas at the time of a scram, the temperature will attain nearly 900 C and a significant thermal stress will occur, which causes severe creepfatigue damage if this occurrence is repeated. It is noted that the primary loads can be disregarded due to no differential pressure since the sleeves are equipped with ventilation gaps and the weight of the neutron absorber is small. The annular space of each absorber section contains a neutron absorber, which is made up of five sintered compacts of B4C/C. The density of the neutron absorber is 1.9 g/cm3.

Control rod support cable

Stand pipe closure Electromagnetic clutch Position indication mechanism Ball screw

AC motor Metal support frame Cable drum AC motor

Neutron absorber section Support ring Neutron absorber

Reserve shutdown hopper Boron -carbide graphite pellet

Control rod guide tube Electric plug

Control rod support cable Reactor pressure vessel

Neutron absorber section Sleeve Neutron absorber spine Shock absorber

Inside sleeve Sleeve Outside sleeve Spine Guide ring Shock absorber A in detail

Figure 2.61 Schematic view of control rod system of HTTR [60].

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2.9.2.2.2 Control rod drive mechanism The control rod drive mechanism consists of an AC motor, a decelerator, an electromagnetic clutch, a velocity-limiting brake, a shock-absorbing mechanism, position indicators, manual control mechanisms, gears, a control rod support cable, etc. as shown in Fig. 2.61. The drive motor is coupled to the cable drum through the decelerator, the electromagnetic clutch, and the gears. During normal operation, the position of the control rod is sustained by the torque of the motor. The maximum withdrawal velocity is limited to a value below 70 mm/s by the decelerator even when the motor is running at the maximum speed even if the motor is uncontrollable. When the electric current through the electromagnetic clutch is cut off by a scram signal, the clutch is separated to insert the control rod into the core by gravity as mentioned before. To protect the control rod support cable from overloading, the velocity-limiting brake maintains the insertion speed constant during a reactor scram by applying the braking torque in proportion to the speed. The shockabsorbing mechanism reduces the speed of the control rods at the last stage of the insertion by using a gas damper to absorb any impact when the control rods are inserted for more than 80% of their full stroke. The position indicator has three independent systems using synchro-mesh transmitters as a built-in safety factor.

2.9.2.3 High temperature structural design guideline of control rod Fig. 2.62 shows the flow diagram for the elevated temperature design of the HTTR control rod [62]. The high temperature design is based on the ASME Boiler and Operation condition category

Load-controlled stress limits

Strain and deformation limits Inelastic Elastic analysis analysis Material strain limits

I and II

TEST No. 1 No. 2 No. 3 No. 4

III

1% 2%

Membrane Bending

Creep fatigue eveluation IV

or When temp. exceeds 900◦C

Prove by test that reactor shutdown function of control rod is maintained

Replace control rod

Figure 2.62 Schematic flow diagram for high temperature design of HTTR control rod [60].

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Pressure Vessel Code Case N-47-21 [28] because the failure modes assumed in the code case are most generic and their integrity evaluation methods are applicable to the design of HTTR control rods. With respect to the factors of the stress limits on Sm: the allowable stress intensity values, laid down in the Japanese structural guideline for core support structures of LWRs are adopted. One notable merit of the HTTR control rod design is that the control rods are replaceable, which makes the structural design easier and lessens the requirements for a database on material properties at the same time. As also shown in Fig. 2.62, the design code for the control rod requires that the control rod be replaced when it has been subjected to temperatures above 900 C. The threshold temperature of 900 C is provided not only to reduce the need for very high temperature material data but also to have a reasonably long life for the control rods [13]. Damage such as cracks to the structural metal is essentially acceptable as far as the reactor shutdown function is secured. In some events of Japanese operation condition categories IIIV (corresponding to levels BD in N47, respectively), the temperature of the control rod can exceed 900 C. In these cases, it must be proved by analysis or test that the reactor shutdown function of the control rods is maintained and the control rods need to be replaced. The analytical approach is possible up to 1000 C. For the design by test, verification tests using a full-size model of the neutron absorber section were carried out above 900 C (at 1100 C maximum), which will be described in Section 2.9.2.5. Significant deformation or cracks were not observed, which allow the control rod to function properly. Thus design of the control rod below 900 C and above 900 C is conducted by “design by analysis” and “design by analysis or test,” respectively. Alloy 800H is selected for the metal parts of the control rod mainly because iron base alloys are superior to nickel-rich alloys in both postirradiation tensile and creep properties [63]. Design material data on Alloy 800H were determined in this study, which covers temperatures of up to 900 C for Smt, St and design fatigue curves, and up to temperature of 1000 C for Sm, Su, and SR, which will be presented in the following section. During normal operation of the high temperature test operation mode, that is, when the control rods are withdrawn from the core, the maximum temperature of the sleeve at the bottom neutron absorber section is approximately 550 C by preliminary temperature analysis [61]. Since the maximum temperature of graphite blocks in the fuel region is calculated to be around 1100 C, temperatures of the control rods in the fuel region exceed 900 C when all the control rods are inserted into the core at the same time at a scram. Then, the inner seven pairs of control rods in the fuel region must be replaced under the design guideline. Thus it is indispensable for the HTTR to employ a two-step control rod insertion method to avoid a control rod replacement. The outer nine pairs of control rods in the replaceable reflector region are inserted into the core first. The maximum temperature of the outer control rods then attains a temperature somewhat lower than 900 C, which will be shown in Section 2.9.2.6. The other inner seven pairs in the fuel region are inserted 40 min later or when the outlet coolant temperature becomes less than 750 C so that the maximum temperature of the inner control rods stays lower than the temperatures of the outer control rods. Out of

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these outer nine pairs of control rods, six pairs, which are positioned nearer to the center of the core than the other three, are subjected to the severest thermal conditions. Stress analysis of one of these six pairs of control rods in the replaceable reflector region was conducted and will be shown in Section 2.9.2.7.

2.9.2.4 Design material data on Alloy 800H The high temperature design of the HTTR control rod is based on the ASME Boiler and Pressure Vessel Code Case N-47-21 as shown in the previous section; however, design material data on Alloy 800H available in the Code Case are below 1400 F (760 C). Allowable stresses and design curves on Alloy 800H, which are needed for the design of control rods, were determined up to temperature of 1000 C in this study based on existing data. Material data in the previous papers [6467] are used for the determination. In the following, design curves of allowable stress intensity values: Sm, Su, St, expected minimum stress-rupture values, denoted SR in this paper, and design fatigue strain range εt are shown. Definitions of the values are the same as those in the Code Case N-47-21. Fig. 2.63AC shows design curves of Sm and Su, St, and SR, respectively. In Fig. 2.63D, relation between shifted curves of test results by Furukawa et al. [68] and design fatigue strain range εt of Alloy 800H at 1400 F (760 C) in the Code Case N-47-21 is depicted. In the figure, the original regression curve in Furukawa

Figure 2.63 Design properties of Alloy 800H: (A) Sm and Su; (B) St; (C) stress-to-rupture; (D) design fatigue strain range [60].

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et al. [67] is shifted by less than 1/20 in cyclic life and 1/2 in total strain range to derive the design fatigue curve. The test results between 850 C and 950 C in Fig. 2.63D indicate very small temperature dependence. The shifted curves at 850 C and 900 C locate at a region of longer life than the design curve of N-47-21 at 1400 F (760 C). At 950 C, the shifted curve and the design curve lie at almost the same position. Thus it is concluded that design fatigue strain range εt of Alloy 800H at 1400 F (760 C) in the Code Case N-47-21 is applicable to the design of the HTTR control rod at 900 C.

2.9.2.5 Results of R&D Various tests were performed to evaluate the reliability of the control rod system. It is assumed in design of the HTTR that the number of scrams during plant lifetime is 211 including ten and one times for service loadings of levels III (C in ASME) and IV (D in ASME), respectively.

2.9.2.5.1 Scram tests of the control rod system under seismic conditions Fig. 2.64 shows the scram test equipment, which consists of a control rod drive mechanism, a pair of control rods, and a control rod guide column at the center of a vibrating table. The whole arrangement is supported by a rigid concrete test tank.

Standpipe

Support vessel

Control rod drive Radiation shield Guide tube Radiation shield block Vibration table

Control rod

Control rod guide block

Block to block gap adjuster Support vessel

Test pit Plenum block 0.95m

Figure 2.64 Schematic view of scram test equipment under seismic condition [60].

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Fixed conditions ① Subtend direction ② Core external surface ③ Maximum block to block gap

Design condition

Design limit

Scram time (s)

Hor.

Hor.+Ver.

No vibration scram time

S1: maximum design earthquake S2: uppermost design earthquake Horizontal excitation acceleration (Gal)

Figure 2.65 Typical results for scram time [60].

Block misalignment and contact of a control rod with the boring wall surface of the block during earthquakes were considered as major causes of disturbance during a scram. The following parameters, therefore, were chosen in the tests: the excitation direction, which is a combination of horizontal and vertical excitations, acceleration, and frequency and block-to-block gaps. Fig. 2.65 shows the typical scram time under seismic conditions. Generally, compared with the design specifications or the scram time during normal operation conditions, the amount of disturbance during a scram was found to be negligible for all test results. Therefore it was concluded that the seismic effect on the ability of scramming the reactor is not significant.

2.9.2.5.2 Reliability test of control rods in the HENDEL loop Tests on the control rod system were performed to verify its reliability with singlechannel test rig of the fuel stack test section in the HENDEL loop under the same operation conditions as anticipated for the HTTR. The control rod drive mechanism was operated reliably far more than the total driving numbers considered for a 20-year HTTR operation. No failure of a scram was observed, that is, the phenomena which make it impossible to insert or withdraw a control rod did not occur, and the control rod drive mechanism could move the control rod to given positions without error. Overruns during a scram were small, ranging from 63 to 71 mm. A pressure drop change in the channel due to the movement of the control rod was found to be so small that it did not affect the flow rate distribution of the coolant in the fuel

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stacks. After the test, parts of the control rod system were disassembled and examined with no damage being observed [68].

2.9.2.5.3 Verification tests of the control rods In addition to the above tests, verification tests using a full-size model of the neutron absorber section were carried out. In the tests, postulated accident conditions and cyclic thermal stress were induced in accelerated conditions by heaters around the test model. The effect of inelastic deformation and creepfatigue damage, induced by the thermal stress, was investigated. As a result, little deformation was seen. In other tests of postulated severe accident conditions, the model was subjected to its own load and to severe temperatures as high as 1100 C. After the tests, the shape of the model was investigated. Its deformation was found to be within the limits that will allow the control rod to function properly.

2.9.2.6 Temperature analysis Preliminary temperature analyses revealed that in several events of operation categories II and III, which induce reactor scrams, the maximum temperature of the control rod can exceed 900 C [61]. The severe events include “Failure in the startup of one of two auxiliary gas circulators,” “Failure in the start-up of one of two auxiliary gas circulators and loss of off-site electric power,” “Leakage from the inner pipe of primary concentric hot gas duct,” and “Leakage from the inner pipe of auxiliary concentric hot gas duct.” At normal operation conditions, the main cooling system of the HTTR removes heat of the reactor. At a scram, the primary gas circulators stop and the auxiliary cooling system starts cooling the reactor core instead. So, trouble of the auxiliary cooling system increases temperature of the control rod. Detailed temperature analyses were performed, which took into consideration improvement of the design of the HTTR control rod such as alteration in dimensions. Among the events of operation categories II and III, the representative event in which temperature of the control rod is predicted to become close to 900 C is “loss of off-site electric power.” This event is due to a failure of the power transmission line for the HTTR electrical equipment. The loss of electric power supply causes all of the gas circulators and pumps to stop. Then the reactor scrams, emergency power feeder starts up, and the auxiliary cooling system starts automatically in less than 60 s after the scram because the emergency electric generator takes 60 s, at the maximum, to supply sufficient electricity. Procedure and result of temperature analysis at normal operation and “loss of off-site electric power” are shown below. The temperature analysis consists of two steps: analysis of the core and analysis of the control rod. Calculations of the two steps were performed using the flow analysis code “TRUMP” [69], which solves problems involving flow in temperature fields, pressure fields, etc. In the first step, 1/12 sector of the HTTR core is threedimensionally modeled. More detailed control rod model is used in the second step as shown in Fig. 2.66. It should be noted that difference of elevation of 100 mm exists between the fuel elements and control rod guide blocks.

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Spine Control rod sleeve

Alloy 800H

Alloy 800H Surface element Upper shield Top replaceable reflectors

Position of control rod at normal operation (lowest position considered for the design)

Position of control rod after a scram (dimension: mm)

1st control rod guide block 1st fuel element 2nd control rod guide block 2nd fuel element 3rd control rod guide block 3rd fuel element 4th control rod guide block 4th fuel element 5th control rod guide block 5th fuel element Bottom replaceable reflectors

Figure 2.66 Model of control rod for flow network analysis [60].

In the thermal design of the HTTR control rod, systematic and random factors are considered to account for various uncertainties [13]. The systematic factors, which are a direct accumulation of conservatism, include total reactor power, primary coolant flow rate, and reactor inlet coolant temperature. The random factors, statistically treated, consist of manufacturing tolerances, uncertainties on physical properties, etc. In temperature analysis for the control rods, systematic factors are multiplied to input data, while temperature increases due to the random factors are added to the calculated result. During normal high temperature test operation, inlet and outlet coolant temperatures of the reactor are controlled within 395 C 6 2 C and 950 C 6 17 C, respectively. The position of the control rod considered in the design at the normal operation is shown in Fig. 2.66, which is the lowest position during the life of the control rod. As it can be easily seen, temperature of the control rod becomes highest at the bottom end of the center control rod, which is nearest to the center of the core. Fig. 2.67 shows results of temperature analyses of the center control rod at the

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Figure 2.67 Temperature of the center control rod at high temperature test operation [60].

normal high temperature test operation. The maximum temperature of the control rod sleeve reaches 600 C. Temperature of the neutron absorber B4C/C is higher than that of the sleeve because of γ-heating of the absorber. In the event of “loss of off-site electric power,” when the scram signal triggers the rods, the outer nine pairs of the control rods in the replaceable reflector region are inserted into the core first. The two-step control rod insertion method is employed. Then the auxiliary cooling system starts up in 60 s to remove residual heat, and flow rate of the reactor inlet/outlet coolant becomes twelve percent of that at normal operation. Fig. 2.68 shows the transient temperature of the reactor inlet and outlet coolant in the event, which was derived from simulations using a model of the whole HTTR system with the THYDE-HTGR code [70]. It is 2400 s (40 min) after the scram that the inner six pairs of control rods in the fuel region are inserted, which allows a further decrease in reactor inlet coolant temperature. Fig. 2.69 shows how temperatures of the control rod sleeve and spine in the replaceable reflector regions change as time elapses in the event of “loss of off-site electric power.” Because the control rod is inserted into the hot well at a scram, the temperature increases rapidly at first, reaches 884 C maximum for the outside sleeve, and then decreases slowly. The systematic factors described above are included in the calculations shown in Fig. 2.69. Considering the random factors, the maximum temperature of the sleeve becomes 898 C, which is less than the design limit of 900 C.

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Figure 2.68 Transient behavior at “loss of off-site electric power” [60].

Figure 2.69 Temperature change of control rod in replaceable reflector region at “loss of off-site electric power” [60].

2.9.2.7 Stress analysis Elastic stress analyses were performed utilizing the well-known FEM code ABAQUS. Primary stress is the largest at the top of the spine, nevertheless, the value is quite small (about 13 MPa). Because temperature change of the sleeve is rapid and repeated, evaluation of its secondary stress, and therefore, creepfatigue damage, is the most important in the design of the control rod. The secondary stress range, QR in the event of “loss of off-site electric power” is calculated to be a maximum of 140 MPa. Fatigue

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damage, Df per event is about 0.008 and creep damage, Dc is 0.002, that is, creepfatigue damage per scram is about 0.01. Because the number of scrams postulated in the HTTR is about 50 for 5 years, which is the target life of the HTTR control rod, cumulated creepfatigue damage is about 0.5. Since the design limit is 1.0, the target life of the control rods of 5 years can be readily achieved [71].

2.9.3 Reserve shutdown system 2.9.3.1 Design Fig. 2.70 and Table 2.25 show an outline and specification of the reserve shutdown system. The reserve shutdown system consists of B4C/C pellets, hoppers that

Position indicator Ball screw

Drive motor

Radiation shield

Reactor pressure vessel

Sleeve Control rod Neutron absorber

Hopper

Plug Guide tube

Figure 2.70 Schematic view of reserved shutdown system [60]. Table 2.25 Specifications of reserve shutdown system [60]. Items

Properties

Driver method

Drop B4C/C pellets by gravity

Number

16 (15 in irradiation tests using the center column)

B4C/C pellets Diameter (mm)

13

Length (mm)

13

Material

Sintered pellets of B4C/C

Height of inserted pellets (10-3m-3)

Outflow volume of pellets (10-3m3)

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(A)

Time (s)

127

Core height Equivalent to 2.1 10-2m3 in volume of pellets Hopper release

(B)

Time (s)

Figure 2.71 Outflow property of pellets (A) Change of outflow of dummy pellets after pulling out the electric plug; (B) Change of stack height of inserted pellets [60].

contain the pellets, driving mechanisms, guide tubes, etc. In accidents when the control rods cannot be inserted, the electric plug is pulled out by the drive motor, and the neutron-absorbing pellets fall into the core by gravity. The reserve shutdown system shall be designed so that the reactor should be made and held subcriticality from any operation condition at temperatures from 27 C (300 K) to 950 C by dropping the pellets in.

2.9.3.2 Results of R&D R&D was carried out to investigate the reliability of the reserve shutdown system, especially its mechanism. Fig. 2.71A shows the relation between the outflow volume of the dummy pellets and the time after an operator begins pulling out the electric plug. The pellets fall at a constant rate after the flow has become stable. Fig. 2.71B shows the stack height of the inserted pellets versus time, calculated from the results of Fig. 2.71A. The pellet stack reaches the height of the core in about 22 s. It was not observed that any of the pellets stuck preventing them from falling into the core. Thus it was confirmed that the reserve shutdown system is able to insert pellets reliably and stably.

2.10

Instrumentation and control system

2.10.1 Introduction The instrumentation and control system consists of the instrumentation, control equipment, and safety protection system. There are not many differences in the instrumentation and control equipment design between the HTTR and light water reactors except some features. Various R&D for reactor instrumentation were performed taking into account the HTTR operational conditions. For the HTTR, some

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detectors were developed in order to measure the neutron flux and the in-core temperature because these detectors have to operate in a high temperature environment in comparison with light water reactors. A detection system of fuel failure was developed because the FP from the failed fuel to the primary cooling system are far smaller than in light water reactors during normal operation [13]. In the HTTR, the temperature of the moderator and the coolant changes slowly with reactor power since the heat capacity of the core is very large. A plant dynamic analysis was carried out considering the operational conditions of the HTTR in order to design the control system of the HTTR [72].

2.10.2 Instrumentation system The instrumentation system consists of reactor and process instrumentations to provide information for operation and reactor protection.

2.10.2.1 Reactor instrumentation The reactor instrumentation monitors the major parameters in the operation condition of the HTTR, such as the neutron flux, the position of control rods, the differential pressure in the core, the coolant temperature at the hot plenum, and FPs from failed fuel.

2.10.2.1.1 Nuclear instrumentation Two types of neutron detectors were developed for the HTTR: One is a fission counter, which is prepared for the wide range monitoring system and is used under a high temperature environment at the top of the permanent reflector; the other is an uncompensated ionization chamber, which is prepared for the power range monitoring system and can detect a low neutron flux level outside the RPV. The wide range monitoring system and power range monitoring system are used in the power range from 1028% to 30% and 0.1% to 120%, respectively. The temperature around the wide range detector becomes about 600 C and the neutron flux level around the power range detector becomes about 107 n/cm2 during rated power operation of 30 MW. Fig. 2.72 shows the arrangement of neutron detector for the wide range monitoring system and power range monitoring system. The neutron detector for the wide range monitoring system is required to be able to detect the neutron flux at 400 C and 600 C for normal operation and design basis accident, respectively. A neutron detector, which can be used at high temperature environment of 600 C, was developed. Though an accelerated irradiation test at 600 C, a long-term in-core operation test at 600 C for 1000 days, and an overheating test at 800 C for about 500 h in simulating the condition of an accident, it was found that this detector could withstand the test. The neutron detector for the power range monitoring system is required to have high sensitivity because the neutron detector is arranged outside the RPV and the neutron flux is about 107 n/cm2s, which is lower than that of light water reactors by the magnitude of two orders. The high sensitivity neutron detector, in which 3He gas is charged, was developed. The sensitivity of the neutron detector for the power

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Figure 2.72 Neutron detector arrangement [13].

range monitoring system was 4.7 3 10212 A/nv and no noticeable change of the output linearity is observed after the irradiation test. The results satisfied the requirement for the nuclear instrumentation of the HTTR.

2.10.2.1.2 Control rods position instrumentation The control rods position instrumentation monitors the position of 16 pairs of control rods. The position of control rods is measured by the encoder sensor in the control rod drive mechanism and the signal from this instrumentation is used for the reactor control system and the safety protection system.

2.10.2.1.3 Three core differential pressure instrumentation The core differential pressure instrumentation detects a decrease in primary coolant flow in the reactor core. The core differential pressure instrumentation measures the differential pressure between the inlet and outlet of the core, and the signal from this instrumentation is used for the safety protection system and is transferred to the central control room.

2.10.2.1.4 Fuel failure detection system It is very important to prevent FPs from being released abnormally into the primary cooling system during the normal operation. The fuel failure detection system detects the failure of CFPs by detecting short-life FPs, such as 88Kr and 138Xe. The conceptual block diagram is shown in Fig. 2.73. The fuel failure detection system is composed of two precipitators, a preamp and a compressor. The schematic drawing of precipitator is shown in Fig. 2.74 [73]. The helium gas of primary cooling system from the seven regions in the hot plenum is transferred to the precipitator chambers.

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Figure 2.73 Schematic view of fuel failure detection system [13].

Figure 2.74 Schematic drawing of precipitator [73].

And FPs in the helium gas are collected by precipitating wire in the precipitator chambers. Collected FPs are transferred to the front of the scintillation detector by wire. Then a scintillation detector detects β-rays radiated from short-life gaseous FPs. In the HTTR, the initial failure shall be less than 0.2% in terms of the sum of heavy metal contamination and SiC defects. The value of 0.2% was determined from the viewpoint of limit of off-site exposure during normal operation. The fuel failure detection system is able to detect 0.002% fuel failure.

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The detection of fuel failure is difficult in comparison with that for conventional metal cladding fuel. A CFP consists of a microsphere of low enrichment UO2 with the TRISO (TRi-ISOtropic) coating. The CFPs are incorporated into fuel compacts with a graphite matrix. If the CFP is failed, the FPs from the failed fuel to the primary cooling system are very small. Furthermore, the radiation background changes depending on the condition of the reactor (neutron flux and fuel temperature and others). To solve these problems, a high-sensitive fuel failure detection system has been developed based on the following considerations: 1. To deduce the effect of long-life FPs accumulated in the primary cooling system and to detect the FPs are detected selectively. 2. The state equation, which identifies the concentration of short-life FPs in the primary cooling system, was established. This state equation was introduced into fuel failure detection system to distinguish the small change in short-life FPs concentration caused by the change in the reactor operation parameter, such as fuel temperature and local power density.

The experiment of fuel failure detection system was performed since the end of 1981 at Japan Material Test Reactor (JMTR). And these results satisfy the requirement to identify 1024% of the fuel damage.

2.10.2.1.5 In-core temperature monitoring system Four thermocouples are arranged at each hot plenum block in order to monitor the primary coolant temperature. The maximum coolant temperature around the thermocouples reaches about 1100 C. N-type thermocouples (NicrosilNisil) are used because the deviation of thermal electromotive force is small compared with other types of thermocouples under a high temperature environment. The thermocouples for the in-core temperature monitoring system should have a long lifetime with high reliability. Stability of thermos-electromotive force of many thermocouples were checked at long-term performance test, which was performed under the condition that the temperature and its running time were 1200 C and 20,000 h, respectively. As a result of the test, N-type thermocouple was selected.

2.10.3 Process instrumentation The process instrumentations of temperature, pressure, flow rate, radioactivity, etc. are required to monitor the plant parameter during the reactor operation. There are about 4000 sensors in the HTTR, and the signals from the sensors are centralized by the plant computer. The process instrumentation is used to measure process parameters in the primary cooling system, secondary helium cooling system, pressurized water-cooling system, etc. The signals of process instrumentation are transferred to the safety protection system, reactor control system, and others. The process instrumentations used for the reactor protection system and the engineering safety features actuating system consist of three identical channels.

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2.10.4 Control system The control system of the HTTR consists of an operational mode selector, a reactor power control system, and a plant control system. Microcomputers are used for the plant control system and the reactor power control system. An operational mode selector supervises them.

2.10.4.1 Operational mode selector The HTTR is designed to achieve several operational modes to perform rated operation, high temperature test operation, safety demonstration test, and others. The operational mode selector is used to specify the control system and the reactor scram point of instrumentation. The reactor has several operational modes. The reactor outlet coolant temperature is set at 850 C or 950 C at the rated power. “Rated operation” stands for the operation at 850 C and “high temperature test operation” stands for the operation at 950 C. The reactor has two loop operational modes, one is called the “parallel-loaded operation” and the other one is the “single-loaded operation.” The IHX and PPWC are operated simultaneously during the former mode. On the other hand, only the PPWC is operated to remove the heat of 30 MW during the single-loaded operation. The maximum heat capacity of the IHX and PPWC is 10 and 20 MW, respectively, during parallel-loaded operation. The HTTR has several safety demonstration test operational modes, one is the “primary coolant flow rate decreasing test mode” and the other one is the “control rod withdrawing test.” The operational mode selector has select switches for selecting the operational modes. The demand values of the control system and the set point of the safety protection system are changed automatically according to the operational mode selector.

2.10.4.2 Reactor power control device The reactor power control device consists of a reactor power control system and a reactor outlet coolant temperature control system. The reactor power and reactor outlet coolant temperature control systems are cascade connected: the latter is an upper control system to give demand to the power control system.

2.10.4.2.1 Reactor power control system The signals from each channel of power range monitoring system are transferred to three controllers using microprocessors. In the case that there is a deviation between the process and set values, a pair of control rods is inserted or withdrawn at control rods speed from 1 to 10 mm/s according to the deviation. The relative position of 13 pairs of control rods, except for three pairs of control rods used only for the scram, are controlled within 20 mm by the control rods pattern interlock to prevent any abnormal power distribution.

2.10.4.2.2 Reactor outlet coolant temperature control system The reactor outlet coolant temperature control system is used at about 100% of the rated power. In the case that there is a deviation, this control system gives a demand

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to the power control system and changes the reactor outlet coolant temperature by moving the position of the control rods.

2.10.4.3 Plant control device The plant control device controls the plant parameters, such as the reactor inlet coolant temperature, the primary coolant pressure, and the differential pressure between the primary cooling system and the pressurized water-cooling system or secondary helium cooling system. The schematic diagram of the plant control device is shown in Fig. 2.75.

2.10.4.3.1 Reactor inlet coolant temperature control system The reactor inlet coolant temperature control system is used in the power range from 30% to 100% and is cascade connected with a pressurized water temperature control system. In the case that there is a deviation, the reactor inlet coolant temperature is controlled by adjusting the pressurized water coolant inlet temperature of the pressurized water.

2.10.4.3.2 Intermediate heat exchanger primary coolant flow rate control system The IHX primary coolant flow rate control system controls the primary coolant flow rate in the IHX at constant value by adjusting the helium gas circulator.

Figure 2.75 Schematic diagram of control system of HTTR [13].

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2.10.4.3.3 Primary-pressurized water cooler primary coolant flow rate control system The PPWC primary coolant flow rate control system controls the primary coolant flow rate in the PPWC at the constant value by adjusting the revolution of the three helium gas circulators.

2.10.4.3.4 Primary helium pressure control system The primary helium pressure control system controls the primary helium pressure by activating the valves of the helium storage and supply system for primary cooling system. This system is used at about 100% of the rated power.

2.10.4.3.5 Primarysecondary helium differential pressure control system The primarysecondary helium differential pressure control system controls the differential pressure between primary and secondary helium by activating the valves of the helium storage and supply system for secondary helium cooling system. The secondary helium pressure is controlled higher than that of the primary by this system so as to prevent any release of FPs into the secondary helium cooling system.

2.10.4.3.6 Primary-pressurized water differential pressure control system The primary-pressurized water differential pressure control system controls the differential pressure between primary helium and pressurized water by activating the valves of the pressurizer in the pressurized water-cooing system. The pressurized water pressure is controlled lower than that of the primary helium by this system so as to prevent any water from entering into the primary cooling system.

2.10.4.3.7 Pressurized water temperature control system The pressurized water temperature control system controls the inlet pressurized water temperature of PWC by adjusting the flow rate of the pressurized water in the air cooler. The flow rate into the air cooler is adjusted by a bypass flow control valve and an outlet system is given by the reactor inlet coolant temperature control system, which is the upper control system in the cascade.

2.10.5 Safety protection system The safety protection system consists of the reactor protection system and engineered safety features actuating system. The reactor protection system ensures the integrity of the core and reactor coolant pressure boundary under abnormal operational conditions. The engineered safety features actuating system prevents FPs from being released into the environment due to an accident, such as a rupture of the primary concentric hot gas duct.

2.10.5.1 Reactor protection system The reactor protection system inserts the control rods into the core to ensure the integrity of fuel and protect the reactor coolant pressure boundary under abnormal operating conditions. The logic circuits of this system have two trains, which

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receive the signals from the reactor and process instrumentation, and send the signals in case of reactor scram. The logic trains of reactor protection system are set with two parallel systems and each logic train is connected to the circuit breaker with two series. Each logic train sends a signal in case of a reactor scram independently. In case of a reactor scram except for a depressurization accident, the control rods are inserted into the replaceable reflector region at first, and then the remaining control rods are inserted into core when the reactor outlet coolant temperature decreased to about 750 C or 40 min has elapsed after reactor scram. In the case of depressurization accident, all control rods are inserted into the core simultaneously.

2.10.5.2 Engineered safety features actuating system The engineered safety features actuating system sends the signals actuating the engineered safety features, such as the isolation valves of containment vessel, the auxiliary cooling system, and the emergency air purification system. The engineered safety features protect the reactor, the reactor coolant pressure boundary, and the containment vessel boundary and prevent large amounts of FPs from being released outside the reactor facility. This system consists of logic circuits having two trains, which receive the signals from the reactor and process instrumentation, and actuates the engineered safety features.

2.10.5.2.1 Signal isolating containment vessel The signal isolating the containment vessel activates to close the isolation value of the containment vessel in order to prevent FPs releases in a depressurization accident. These signals also activate to stop the air supply and exhaust in the ventilator and air conditioner in the reactor building system and to start up the emergency air purification system.

2.10.5.2.2 Signal starting up auxiliary cooling system The signals starting up the auxiliary cooling system, which are the signals for a reactor scram, activate to start up the auxiliary cooling system so as to remove residual heat in case of a reactor scram except for the case of the depressurization accident or an auxiliary heat exchanger heat transfer tube rupture accident.

2.10.5.2.3 Signal isolating auxiliary cooling water line The signals isolating the auxiliary cooling water line activate to stop the auxiliary cooling system and to close the valves of the containment vessel connected to the auxiliary heat exchanger and the valve of the primary coolant pressure boundary connected to the primary helium purification system.

2.10.6 Performance test results In the “rise-to-power test,” many performance tests were carried out in order to confirm the characteristics of the instrumentation and control system of the reactor.

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2.10.6.1 Characteristics of the neutron flux monitoring system The wide range monitoring system and the power range monitoring system are used in the power range from 1028% to 30% and 0.1% to 120% of the rated power, respectively. The measurement range of the power range monitoring system should overlap the measurement range of the wide range monitoring system in order to continuously monitor the neutron flux. Therefore it is very important to confirm that the measurement range of the wide range monitoring system and the power range monitoring system overlap. Fig. 2.76 shows the calibrated result of the wide range monitoring system and the power range monitoring system from 0.1% to 30% of the rated power. It can be found that response values of the power range monitoring system are nearly equal to those of the wide range monitoring system within the reactor power range from 1% to 10%. This result satisfies the design requirement concerning the overlap between the wide range monitoring system and the power range monitoring system.

2.10.6.2 Primary-pressurized water cooler primary coolant flow rate control system The PPWC primary coolant flow rate control system controls the primary coolant flow rate in the PPWC at a constant value by adjusting the revolution of three helium gas circulators. This control system has been designed as P-I control system with delayed time, and is required to be stably controlled for a 6 10% stepwise change. In order to confirm the stability of this control system, 210% stepwise change from 15.1 to 13.6 t/h was set at the set value of the primary coolant flow rate. Fig. 2.77 shows the test result. The primary coolant flow rate gradually decreased, and reached

Figure 2.76 Calibrated result of wide range monitoring and power range monitoring systems [73].

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137

13.6 t/h without undershoot. The test result shows that the PPWC primary coolant flow rate control system has stable response characteristics against disturbances.

2.10.6.3 Reactor outlet coolant temperature control system The reactor outlet coolant temperature control system gives a signal to the power control system and controls the reactor outlet coolant temperature by controlling the position of the control rods. The response test for this control system has been carried out to confirm the behavior against ramp-wise change under actual operational condition. Fig. 2.78 shows the

Figure 2.77 Test result of primary-pressurized water cooler primary coolant flow rate control system [73].

Figure 2.78 Test result of the reactor outlet coolant temperature control system under the ramp-wise change from 800 C to 812 C [73].

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High Temperature Gas-cooled Reactors

transient behavior for a 115 C/h ramp-wise change from 800 C to 812 C. The reactor outlet coolant temperature control system gives a demand for a reactor power increase to the reactor power control system. The set value of reactor power is increased up gradually from 95% to 98% at the rated power according to the deviation between the reactor outlet temperature and the set value of reactor outlet temperature. In proportion to the rise in the set value the reactor power rises gradually by withdrawing the control rods. Then, the reactor outlet coolant temperature rises gradually with the reactor power rise. The reactor outlet coolant temperature has been reached at 812 C stably. The test result shows that the reactor outlet coolant temperature control system has the capability of controlling the reactor outlet temperature stably.

2.11

Containment structures

2.11.1 Introduction The containment structures of the HTTR consist of the reactor containment vessel, the service area, and the emergency air purification system, which minimize the release of FPs during accidents with FP release from the reactor facilities. The reactor containment vessel is designed to withstand the temperature and the pressure transients and to be leak-tight within the specified limits in the case of a rupture of the primary concentric hot gas duct (depressurization accident). The pressure inside the service area will be maintained at a negative pressure by the emergency air purification system. Radioactive materials are released from the stack to the environment via the emergency air purification system during accident conditions. The emergency air purification system will remove airborne radioactivities and will maintain the design pressure in the service area [74].

2.11.2 Reactor containment vessel 2.11.2.1 Design and construction The reactor containment vessel is comparatively small to minimize the amount of air, which may react with graphite components in the event of a rupture of the primary pressure boundary. The reactor containment vessel is made of carbon steel 30.3 m in height, with 18.5 m inner diameter, and 2800 m3 free volume. Its configuration is shown in Fig. 2.79. The personal air lock, which has an inner door and outer door, is provided for entrance into the reactor containment vessel. The maintenance hatch is equipped for carrying tools and machineries for annual inspections. Elastic plugging material is installed between the reactor containment vessel and base mat of the reactor building to absorb any thermal expansion in case of accidents [75]. Fig. 2.80 shows the sketch of the reactor containment vessel with its penetrations. The reactor containment vessel is made to be pressure-proof during depressurization accidents and the pressure-proof test is performed 1.125 times its maximum pressure of 0.4 MPa. The reactor containment vessel is also made to be leak-tight

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Figure 2.79 Schematic view of reactor containment vessel and its attached equipment [75].

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High Temperature Gas-cooled Reactors

Figure 2.80 Schematic view of reactor containment vessel with its penetrations [75].

and its leakage rate is designed to be less than 0.1% of the total free volume of 2800 m3 per day at room temperature, air atmosphere, and 0.9 times its maximum pressure (0.36 MPa). During a depressurization accident, the peak pressure is 0.36 MPa, which appears 9 s after the initiation of an accident, and the temperature rises to 85 C after 1 s [13]. In the safety case, a leakage rate of 0.25 %/day was used for the calculation to simulate the accident condition as the increase of temperature and pressure and the difference between helium and air. The specification of the reactor containment vessel is listed in Table 2.26. The construction of the reactor containment vessel started in 1991 at the factory. After various kinds of inspections, including material inspections, such as fracture toughness tests, had performed the reactor containment vessel was transported by parts using ships to the HTTR site. At the HTTR site, performance tests were conducted in order to confirm its initial pressure-proof and leakage characteristics. Fig. 2.81 shows a photo of the reactor containment vessel taken during its initial pressure-proof test in November 1992. After the whole body was constructed, penetrations were equipped with their isolation valves and the reactor containment vessel was completed in 1996 [75].

2.11.2.2 Leakage-rate test In Japanese LWRs, the leakage-rate test of the reactor containment vessel was performed with an open reactor coolant pressure boundary to simulate an accident. In HTGRs, the leakage-rate test is planned with a closed reactor coolant pressure boundary to avoid the release of FPs into the environmental of the reactor containment vessel. It is necessary to establish other test method in order to estimate the leakage rate of the reactor containment vessel with closed reactor coolant pressure boundary. In the HTTR, the leakage-rate test was performed with a closed reactor coolant pressure boundary and additional corresponding tests were conducted. First is the

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Table 2.26 Specification of reactor containment vessel [74]. Containment type

Steel containment

Maximum pressure in service (MPa) 

Maximum temperature in service ( C)

0.4 150

Major size Inner diameter (m)

18.5

Overall height (m)

30.3

Body thickness (mm)

30

Upper head closure thickness (mm)

38

Refueling hatch diameter (m)

8.5

Maintenance hatch diameter (m)

2.4

Personal air lock diameter (m)

2.5

3

Free volume (m )

2800

Material

Carbon steel

Leakage rate

Less than 0.1 %/day at the room temperature and 0.9 times the maximum pressure of 0.4 MPa

Figure 2.81 Photo of reactor containment vessel (30.3 m in height, 18.5 m in inner diameter, and 2800 m3 of free volume) taken during initial pressure-proof test in November 1992 [74].

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High Temperature Gas-cooled Reactors

Figure 2.82 System configuration of whole leakage-rate test of reactor containment vessel [75].

partial leakage-rate test. The partial leakage-rate tests were inspected at penetrations and isolation valves connected to the primary coolant pressure boundary. In an accident, such as a rupture of the hot gas duct, the penetrations and isolation valves connected to the primary coolant pressure boundary, which are also the reactor containment vessel boundary, will be pressurized. Adding these values of partial leakage rate of penetrations and isolation valves to the whole leakage rate can simulate an accident, such as a rupture of the hot gas duct. Second is the lower set pressure of the reactor coolant pressure boundary. The reactor coolant pressure boundary is kept at a pressure below the test pressure during the whole leakage-rate test of the reactor containment vessel in order to avoid a decrease of the measured value in the whole leakage test. The test chart of the whole leakage-rate test of the reactor containment vessel is shown in Fig. 2.82. The allowance value of leakage rate Ldo is regulated as   Ldo %=day 5 Ld ð1 2 A1 Þ 5 0:1 3 ð1 2 0:1Þ 5 0:09;

(2.1)

where Ld is the leakage rate at the room temperature and 0.9 times its maximum pressure of 0.4 MPa, as shown in Table 2.26, and A1 is the factor of deterioration. The deterioration factor is regulated by the Japanese test standard (JEAC4203), and when the test is planned annually, the deterioration factor is 0.1 and for an inspection every 2 years, it is 0.2. The term “leakage rate” is used with the follow meanings: 1. Partial leakage rate LP is the leakage of penetrations and isolation valves connected to the reactor coolant pressure boundary. 2. Conversion leakage rate LCV is the converted value of the partial leakage rate LP to that of the whole leakage rate.

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3. H hours leakage rate LH is the leakage rate from beginning to H hours elapsed at the whole leakage-rate test. 4. Average leakage rate LAV is the leakage rate of the measured value of the whole leakagerate test with 95% confidence limit as for statistical processing and it is shown such as a 6 b. 5. Leakage rate L is the total value of the leakage rate and it is the sum of LCV and upper value of LAV (a 1 b).

2.11.2.2.1 Partial leakage-rate test The partial leakage-rate tests are performed for the penetrations and isolation valves of the primary helium purification system, primary helium sampling system, and fuel failure detection system, which are connected to the reactor coolant pressure boundary. The partial leakage-rate test was executed in accordance with the pressure drop method regulated by the Japanese test standard (JEAC4203). The specification of detectors is shown in Table 2.27. The partial leakage-rate test is conducted at the pressure above 0.36 MPa, which is 0.9 times the maximum pressure of 0.4 MPa. The partial leakage rate LP and conversion leakage rate LCV are calculated by the following equations:   24 P1 2 P2 24 ΔP 3 100 5 3 100; LP %=day 5 H P1 H P1

(2.2)

  Vi ; LCV %=day 5 LP 3 V0

(2.3)

where P is the gauge pressure Pa at arbitrary time, V0 is the free volume of the reactor containment vessel (2800 m3), Vi is the free volume of the partial leakage-rate test area m3, H is the elapsed time h, ΔP is the pressure drop (5P1 2 P2) Pa, subscript 1 is the starting value, and subscript 2 is the value after H hours elapsed. The total value of conversion leakage rate ΣLCV was 1.43 3 1024%/day measured in 1996, and the other years’ results had almost same value [76].

2.11.2.2.2 Whole leakage-rate test The whole leakage-rate test was performed for the reactor containment body and its attached equipment, such as personal air lock, refueling hatch, and maintenance hatch. The whole leakage-rate test was executed in accordance with the absolute pressure method regulated by the Japanese test standard (JEAC4203). The specification of detectors is shown in Table 2.28.

Table 2.27 Detectors of partial leakage-rate test for penetrations and isolation valves [1]. Item

Detector

range

Accuracy

Number

Pressure

Pressure gauge

00.59 MPa

6 0.5% of full scale

2

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High Temperature Gas-cooled Reactors

Table 2.28 Detectors of whole leakage-rate test for body and its attached equipment [74]. Item

Detector

Range

Accuracy

Number

Absolute pressure

Quartz manometer

00.67 MPa

6 0.01% of full scale

1

Temperature

Resistance temperature sensor

0 C50 C

6 1 C

28

Humidity

Dew point detector

0.312 kPa

6 0.02% of full scale

7

In the absolute test method, pressure, temperature, and humidity are obtained. The average temperature T and average dew point PC are: T ð C Þ 5

X

Ti Di ;

(2.4)

i

P C ð C Þ 5

X

Pci Ei ;

(2.5)

i

where Ti is individual temperature measured by resistance temperature sensors, Pci is individual dew point, and Di and Ei are the factors which depend on room volume. The amount of leakage Q, which is the value of weight percent (wt.%) between air in the reactor containment vessel and leaked air, and LH, which is the H hours leakage rate %/day, are calculated by:   G1 2 G2 Pm2 T1 5 12 Q ð%Þ 5 3 100; G1 Pm1 T2

(2.6)

  24Q 3 100; LH %=day 5 H

(2.7)

where G is the air weight kg in the reactor containment vessel, Pm is the absolute pressure of air Pa in the reactor containment vessel, which is the difference value of the measured pressure value by the quartz manometer and the measured humidity value by the dew point detector, and T is the temperature K in the reactor containment vessel. The regression line is analyzed for the Eq. (2.6) and its variance analysis is performed in order to clarify the significant difference between the amount of leakage and the elapsed time. The average leakage rate LAV %/day can be calculated with its 95% confidence limit to LH, and LAV is shown as:   (2.8) LAV %=day 5 24ðb 6 2:07σÞ;

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where b is the amount of leakage in 1 h, which is calculated by statistical processing and σ is also the statistic value. Fig. 2.83 shows the transient of pressure and amount of leakage Q measured during the whole leakage-rate test performed in 1999. The pressure inside the reactor containment vessel was raised by air compressors until the pressure reaches its test pressure of 0.36 MPa. After stability condition was confirmed, the measurement of leakage rate was begun. At the measurement duration, the amount of leakage Q was increased by the leak from the reactor containment vessel, and this amount was converted to the leakage rate with statistical processing. The leakage rates were obtained from 0.017 to 0.034 %/day at eight times testing confirmed from 1996 to 2012 as shown in Table 2.29. These values are well below the set limitation of 0.09 %/day, which includes the deterioration factor.

Figure 2.83 Transient of pressure and amount of leakage Q during second whole leakagerate test [74]. Table 2.29 Leakage-rate test results of reactor containment vessel [74]. 1st

2nd

3rd

4th

5th

Time of whole leakage-rate test

Oct. 1996

Sep. 1999

Dec. 2000

Sep. 2001

Dec. 2002

Leakage rate L (%/day)

0.018

0.017

0.034

0.018

0

6th

7th

8th

9th

10th

Time of whole leakage-rate test

Dec. 2003

Dec. 2004

Dec. 2006

Dec. 2000

Mar. 2012

Leakage rate L (%/day)

0

0.015

0

0.029

0.017

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High Temperature Gas-cooled Reactors

2.11.3 Service area 2.11.3.1 Design The service area, as shown in Fig. 2.84, is the space surrounding the reactor containment vessel where the fuel handling and storage systems, the primary helium purification system, etc. are located. The service area is kept at a negative pressure during accidents by an emergency air purification system, which operates under the isolated reactor containment vessel condition. During normal operation, the service area is also kept at a negative pressure by the ventilation and air conditioning system. In operation floor at the ground level TP 36.5 m, the fuel handling and storage systems are equipped.

2.11.3.2 Commissioning tests As a commissioning test, an airtight test of the service area is conducted. During an accident, like a failure of the primary helium purification system located inside the

Figure 2.84 Schematic view of service area [74].

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Table 2.30 Relationship between air pressure in service area and the elapsed time after start-up of emergency air purification regulation damper opening at 55% [74]. Elapsed time (min) 0

1

2

3

4

5

Air pressure in service area at operating emergency air purification system A (Pa)

17

275

286

282

285

287

Air pressure in service area at operating emergency air purification system B (Pa)

18

271

279

281

281

281

service area, the service area is isolated automatically by an isolation signal. The service area maintains its air pressure of less than 259 Pa by the emergency air purification system. In the airtight test of the service area, the opening of air regulation damper of the emergency air purification system changes from 20% to 55% as a parameter. Table 2.30 shows the relationship between air pressure in the service area and elapsed time after the start-up of the emergency air purification system. During the emergency air purification system operating, air pressure of the service area was kept well below its allowable limitation of 259 Pa within 1 min from the start-up of the emergency air purification system.

2.11.4 Emergency air purification system 2.11.4.1 Design The emergency air purification system is an engineered safety feature and has two independent systems. It removes airborne radioactivity and maintains design pressure in the service area of below 259 Pa during accidents. Each independent system is composed of an exhaust filtering unit, exhaust blower, and automatic butterfly valves. The exhaust filtering unit discharges the purified air to atmosphere through an exhaust duct. Table 2.31 shows the major specifications and Fig. 2.85 shows the schematic diagram of the emergency air purification system.

2.11.4.2 Commissioning tests 2.11.4.2.1 Start-up test The emergency air purification system is installed to mitigate the influence of a failure of primary helium purification system in the service area, etc. In the safety case, the start-up time of the emergency air purification system is important to prevent the public from suffering excessive radiation exposure. Considering the radiation exposure conservatively, the start-up time was set to be 13 min. The start-up test was conducted with the following sequence, which demonstrates a failure of primary helium purification system. This failure is initiated by a rupture in the piping. The reactor containment vessel is isolated by closing the

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High Temperature Gas-cooled Reactors

Table 2.31 Major specifications of emergency air purification system [74]. Exhaust filter unit Type

Corpuscle and iodine removal filter

Number

2

Volume velocity

3360 m3/h

Charcoal layer thickness

50 mm

Allowable limit of removal efficiency Corpuscle

More than 99%

Iodine

More than 95%

Figure 2.85 Schematic diagram of emergency air purification system [77].

isolation valves, running down the pumps of the pressurized water-cooling system, etc. after detecting the increase in the primary helium purification flow within 5 s after a rupture. At the same time, the emergency air purification system is actuated. The reactor is automatically scrammed by detecting a signal of “pressurized water flow rate of the primary pressurized water cooler (PPWC) is low”; isolation valves of the reactor containment vessel are closed, and the auxiliary cooling system starts automatically to remove the residual heat of the core. The sequence of this accident is shown in Fig. 2.86.

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Figure 2.86 Sequence of failure of primary helium purification system [74].

For this sequence, the following times were regulated in the safety case: 1. From the start-up command to the start-up of the blowers and electric heaters of the emergency air purification system. 2. From the start-up time of the blowers and electric heaters to the achievement of 4 C temperature difference between inlet and outlet of the electric heaters. 3. From the loss of off-site electric power simulation to the restart of the emergency air purification system by the stand-by power supply. 4. From the restart-up time of the blowers and electric heaters to the achievement of 4 C temperature difference between inlet and outlet of the electric heaters.

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Table 2.32 Start-up test result of emergency air purification system [74]. Elapsed time (min) Regulation value

Measured value System A

System B

1

Time from the start-up command to the start-up of blowers and electric heaters of the emergency air purification system



0.09 s

0.08 s

2

Time from the start-up time of the blowers and electric heaters to the achievement of 4 C and electric heaters of the emergency air purification system

5

1

1

3

Time from the loss of off-site electric power simulation to the restart-up of the emergency air purification system by the stand-by power supply

1

49.8 s

53.8 s

4

Time from the restart up time of the blowers and electric heaters to the achievement of 4 C temperature difference between inlet and outlet of the electric heaters

5

1

1

Total

13a

.3

.3

a

Total time of 13 min is the sum of item 2 (5 min), item 3 (1 min), item 4 (5 min), and a safety margin of 2 min.

The test result is shown in Table 2.32. The start-up of the emergency air purification system was less than 3 min, which was well below the allowable limitation of 13 min.

2.11.4.2.2 Filter efficiency measurement Filter efficiency of the emergency air purification system for removal of radioactive particles and iodine was confirmed. The allowable efficiency limits of the particle removal and iodine removal are more than 99% and 95%, respectively. The concentration of the smoke of dioctyl phthalate (DOP) particles of 0.7 μm, generated by the temporary testing equipment upstream and downstream of the filter, was detected in order to measure the efficiency for corpuscle removal. The iodine removal efficiency was evaluated by the adsorption efficiency of active carbon and the leakage rate of Freon gas (R-112) through a bypass filter. The corpuscle removal efficiency EC% and iodine removal efficiency EI% are calculated by:   Cd E C ð%Þ 5 1 2 3 100; (2.9) Cu

Design of High Temperature Engineering Test Reactor (HTTR)

 Df 2 Da EI ð%Þ 5 A 3 1 2 ; Uf 2 Ua

151



(2.10)

where Cd and Cu are DOP concentrations ppm downstream and upstream of the filter, respectively, A is the adsorption efficiency of active carbon%, which is the inspected value, Uf and Df are the Freon gas concentrations ppm upstream and downstream of the filter, respectively, and Ua and Da are the air concentrations ppm upstream and downstream of the filter. As a result of the filter efficiency test, the corpuscle removal and iodine removal efficiency were estimated to be 99.99% and 99.59%, respectively, which were well over the allowable limitation of 99% and 95%, respectively.

2.12

Other systems

2.12.1 Introduction Chemistry control is important for the helium coolant of high temperature gascooled reactors because impurities cause oxidation of the graphite used in the core and corrosion of high temperature materials used in the heat exchanger, etc. The helium purification systems are installed in the primary and secondary helium cooling systems in order to reduce the quantity of chemical impurities. The helium sampling systems monitor the concentration of impurities. The helium storage and supply systems keep the steady pressure of the helium system during the normal operation [13]. The new and spent fuels should be handled and stored safely and reliably by the fuel handling and storage system. The fuel handling system is utilized to keep the helium boundary without entering air to the helium system during the refueling, which will be performed every 3 reactor years [13]. This section describes the outline of the auxiliary helium and fuel systems.

2.12.2 Auxiliary helium systems 2.12.2.1 Helium purification system The helium purification system is installed in the primary cooling system and secondary in order to reduce the quantity of chemical impurities such as hydrogen, carbon monoxide, water vapor, carbon dioxide, methane, oxygen, and nitrogen. The primary helium purification system is mainly composed of a precharcoal trap, an inlet heater, two copper oxide fixed beds, coolers, two molecular sieve traps, two cold charcoal traps, and gas circulators as shown in Fig. 2.87 [78]. The flow diagram of the secondary helium purification system is almost same as that of the primary system except for the absence of a precharcoal trap.

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High Temperature Gas-cooled Reactors

Figure 2.87 Flow diagram of primary helium purification system [78].

Table 2.33 Main specifications of primary helium purification system [13]. Item

Type

Number

Helium flow rate (kg/h)

Precharcoal trap

Vertical cylinder

1

200

Copper oxide fixed bed

Vertical cylinder

2

200

Molecular sieve trap

Vertical cylinder

2

200

Cold charcoal trap

Vertical cylinder

2

50

The primary helium gas is introduced into the primary helium purification system through the auxiliary cooling system and the purified helium gas returns to the auxiliary cooling system and stand pipes. Main specifications of the primary helium purification system are shown in Table 2.33. The flow rate of the primary helium purification system is determined considering the following requirements: 1. To satisfy the concentration limit of impurity to reduce oxidation of core graphite structures (limited impurity concentration value is shown in Table 2.34). 2. To purify .10% of the helium inventory in the primary cooling system in 1 h. 3. To purge the stand pipes.

The primary helium purification system has three kinds of traps for reducing chemical impurities. Each trap has two identical systems for reliability of plant operation. The first trap is a copper oxide fixed bed where hydrogen and carbon monoxide are oxidized to water vapor and carbon dioxide, respectively. It is kept at a temperature of 280 C during its operation. The second trap is a molecular sieve trap where water vapor and carbon dioxide are removed by adsorption. The third trap is a cold charcoal trap where oxygen, nitrogen, methane, and noble gases are removed by adsorption; it is kept at a temperature of 2195 C. The flow rate in the primary helium purification system is 200 kg/h. The flow rate through the cold charcoal trap

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Table 2.34 Upper impurity limit in primary coolant at 4 MPa and reactor outlet coolant temperature from 800 C to 950 C [78]. Item

Concentration (ppm)

H2

3.0

CO

3.0

H2O

0.2

CO2

0.6

CH4

0.5

N2

0.2

O2

0.04

Figure 2.88 Efficiency of water vapor removal at molecular sieve trap of primary helium purification system [78].

is 50 kg/h with a bypass flow for the rest of the gas. The efficiency of water vapor removal was confirmed in its commissioning test and is shown in Fig. 2.88. When the efficiency of the traps decreases during operation, the traps are changed manually to stand-by traps. The deteriorated traps can be used repeatedly after regeneration. Noble gases, absorbed by a cold charcoal trap, are stored for about 50 days and are then transferred to the gaseous radioactive waste treatment system. The flow rate of the secondary helium purification system is 10 kg/h, which is determined in a similar manner to that of the primary helium purification system. The design of helium purification system takes credit of the experience obtained in the HENDEL loop [79].

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High Temperature Gas-cooled Reactors

2.12.2.2 Helium sampling system The helium sampling systems detect chemical and radioactive impurities in the primary cooling system and secondary helium cooling system. The concentration of chemical impurities, hydrogen, carbon monoxide, water vapor, carbon dioxide, methane, nitrogen, and oxygen is measured by the gas chromatograph mass spectrometers. The primary helium sampling system, consisting of sampling equipment, a carrier gas supply system, and a standard gas supply system, automatically transmits the impurity concentration measurement to the main control room, as does the secondary helium sampling system. The purpose of the primary helium sampling system is: 1. To monitor the chemical impurity level for the purpose of avoiding core graphite oxidation and carbon deposits, as well as the carburizing and decarburizing of Hastelloy XR in the IHX. 2. To detect the rupture of a heat exchanger tube in the PPWC and auxiliary heat exchanger. 3. To monitor the performance of the traps in the primary helium purification system.

The sampling locations of impurities except water vapor for the primary are the inlet and outlet of the reactor, the inlet and outlet of the primary helium purification system, and the inlet and outlet of the cold charcoal trap. Two detectors are installed for detecting water vapor. Sampling locations are as follows: the inlet to the reactor, the inlet and outlet of the primary purification system, and the inlet to the cold charcoal trap for the “No. 1” detector, the reactor outlet, the outlet of the PPWC, the outlet of the primary helium gas circulator for the PPWC and the IHX for “No. 2.” The sources of the initial impurities in the primary coolant can be traced back to the core graphite, to the heat insulator in the primary hot gas duct, and to original impurities of the helium gas primary coolant. The total amount of water removed during the commissioning tests had been 0.75 kg.

2.12.2.3 Helium storage and supply system The helium storage and supply systems are installed for the primary cooling system and secondary helium cooling system. The primary helium storage and supply system is composed of storage tanks, a supply tank, helium compressors, etc. The primary coolant is kept at a fixed pressure at about 4.0 MPa by the primary helium storage and supply system during normal operation. The secondary helium storage and supply system stores and supplies helium gas and controls the pressure of the secondary helium gas to be higher than the primary coolant to aggravate the entry of primary coolant to the secondary coolant during an accident such as a rupture of the boundary between the primary and secondary helium cooling system. The main specifications of the primary helium storage and supply system are shown in Table 2.35. Six storage tanks with a capacity of 220 kg each and a supply tank with a capacity of 110 kg are installed. Two helium gas compressors are also installed in the system. One compressor is a stand-by system. The pressure of the primary helium gas is maintained at the design value by this system. The supply valve for the primary coolant opens and closes at a pressure of 3.92 and 3.99 MPa, respectively, and the exhaust

Design of High Temperature Engineering Test Reactor (HTTR)

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Table 2.35 Main specifications of primary helium storage and supply system [13]. Item

Number

Tank volume (m3/tank)

Capacity (kg)

Max. pressure (MPa)

Storage tank

6

18

1320 (total)

8.6

Supply tank

1

10

110

8.6

valve opens and closes at a pressure of 4.00 and 3.96 MPa, respectively. The supply of helium gas to the primary coolant is also utilized for detecting the leakage rate of the primary coolant during operation when its automatic pressure control system is working. The very small leakage rate of the primary coolant can be detected by measuring the opening time of the helium gas supply valve. The secondary helium storage and supply system is composed of a storage tank with a capacity of 25 kg, a supply tank with a capacity of 10 kg, two compressors, etc. This system is designed to control the pressure of the secondary helium gas to maintain the differential pressure between the primary and the secondary coolant. The supply valve for the secondary coolant opens and closes at a pressure difference of 229.4 and 29.8 kPa, respectively, and the exhaust valve opens and closes at a pressure difference of 29.4 and 19.5 kPa, respectively. The design pressure difference between the primary and the secondary coolant is 73.5 kPa.

2.12.3 Fuel system The new fuel and the spent fuel should be handled and stored safely and reliably by the fuel handling and the fuel storage systems [80]. Each fuel element in the core will be discharged every three reactor years. The fuel elements keep their original position during their lifetime in the core.

2.12.3.1 Fuel handling system The fuel handling system is utilized to install and remove fuel elements, replaceable reflector blocks, top shielding blocks, control rod guide blocks, and control rods. The fuel handling system consists of a fuel handling machine, attached equipment, and auxiliary equipment. The fuel handling machine consists mainly of a shielded cask, a gripper, a fuel handling unit drive system, a rotating rack, and a door valve. The fuel handling machine has a shield sufficient to protect fuel handling personnel and a gas-tight boundary. With the fuel blocks stored to its capacity, the fuel handling machine maintains subcriticality. The door valve, connected with the bottom of the fuel handling machine, has a gas-tight shield structure along with the reactor isolation valve. The attached equipment used during refueling consists of a reactor isolation valve, connecting pipe, and control rod handling machine. The fuel handling procedure is schematically shown in Fig. 2.89. The fuel elements are refueled column-wise through stand pipes by the fuel handling machine, which is able to handle the five fuel elements at one refueling step.

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High Temperature Gas-cooled Reactors

Figure 2.89 Schematic plan of fuel handling procedure (A) Remove of reactor containment vessel. (B) Connection status of control rod handling machine to stand pipe. (C) Replacement of control rod handling machine with fuel handling machine. (D) Structure of gripper for fuel. (E) Insertion of finger of gripper into refueling hole of fuel block. (F) Removal of spent-fuel by lifting the gripper and storage in fuel handling machine. (G) Storage of spent-fuel in spent-fuel storage pool after moving fuel handling machine to pool. (H) Storage of new-fuel in fuel handling machine after moving it to new-fuel storage cell. (I) Installation of new-fuel into reactor with fuel handling machine [13].

Design of High Temperature Engineering Test Reactor (HTTR)

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Prior to the refueling operation, the reactor is shut down and depressurized, and the refueling hatch of the reactor containment vessel is removed (Fig. 2.89A). The connecting pipe, the reactor isolation valve, and control rod handling machine are installed on the stand pipe located above the region being refueled (Fig. 2.89B). The control rod handling machine removes the stand pipe closure and control rod drive mechanism. After removing the control-rod handling machine, the fuelhandling machine is connected to the reactor isolation valve in order to open the gate of reactor isolation valve (Fig. 2.89C). The gripper automatically descends into the RPV and comes to the position of the top shielding block on the fuel elements. Slowly lowered while the fingers of the gripper enter the handling hole in the block, the gripper positioning device is affixed to the machine to insert the gripper into the handling hole smoothly. After the fingers grab the block in the refueling hole, the gripper is lifted, and the block is put into the rotating rack of the machine (Fig. 2.89CF). When all elements are in the rotating rack, the gate of the reactor isolation valve is closed and the fuel handling machine is transferred to the spentfuel storage pool (Fig. 2.89G). Spent-fuel elements in the fuel handling machine are put into the rack in the pool. Then, the fuel handling machine is transferred to the new-fuel storage cell to put new-fuel elements into the rotating rack (Fig. 2.89H). The fuel handling machine is connected again to the reactor isolation valve and installs new-fuel elements into the RPV (Fig. 2.89I). The duration of the whole refueling is estimated about several months.

2.12.3.2 Fuel storage system The HTTR has two fuel storage systems, one is the new-fuel storage system and the other is the spent-fuel storage system. The new-fuel storage system consists of fuel assembling and testing equipment, a new-fuel storage cell, and inert gas replacement equipment. The new-fuel storage cell can store about one and a half core inventories. The storage rack forming a vessel of a vertical cylinder with a plug has sufficient distance to the adjacent storage racks in order to keep subcriticality. The inert gas replacement equipment evacuates air in the rack and replaces it with pure helium gas to keep fuel elements in a dry condition. The functions of the new-fuel storage system are to inspect the new-fuel rods, to assemble the fuel elements, and to store new-fuel elements. The spent-fuel storage system located inside the reactor building, consisting of a spent-fuel storage pool, water-cooling and purification system, and irradiated material storage pit, stores spent-fuel elements, control rod guide blocks, and replaceable reflector blocks. The spent-fuel storage pool has sufficient shielding and can store spent-fuel elements of about two core inventories. The inside of the spent-fuel storage pool is lined with stainless steel to prevent leakage of pool water. Leakage can be detected by monitoring the water from the leakage check ditch located within the lining. After 2 years cooling in the spent-fuel storage-pool, the spent-fuel is transferred to another spent-fuel storage system located neighborhood building of the HTTR.

158

2.13

High Temperature Gas-cooled Reactors

Safety design

2.13.1 Introduction As for the HTTR Licensing, in February 1989, the application for its installation permit was submitted to the prime minister of Japan by JAERI, which is the present JAEA. In November 1990, the safety review was terminated by Nuclear Safety Commission of the time and the prime minister issued the installation permit of the HTTR to JAERI. Up to now, the significant experimental data to validate the safety design of the HTTR have been accumulated through the following tests: 1. Long-term operation of 30 days at 850 C 2. Long-term operation of 50 days at 950 C 3. Safety demonstration tests a. Reactivity insertion test by control rod withdrawal b. Coolant flow reduction test by tripping of gas circulators c. LOFC test

2.13.2 Basic safety design philosophy Fig. 2.90 shows a logical flow to establish the basic safety design philosophy of the HTTR. The top-level regulatory criteria for the HTTR is identical to that of the

Figure 2.90 Logical flow to establish a safety design philosophy of HTTR [81].

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Table 2.36 Summary of top-level regulatory criteria for HTTR [81]. Dose limits

Top-level regulatory criteria

Normal operation

G

G

Accident

No significant risk of radiation exposure to the public

G

Effective dose equivalent shall not exceed 5 mSv

G

G

G

Hypothetical accident

Based on “Examination Guide for Dose Goal outside the site boundary of Light Water Nuclear Reactor Facilities”

G

G

Major accident

5 mSv of annual radiation exposure outside the site boundary

G

G

G

G

Based on “Examination Guide for Safety Evaluation of Light Water Nuclear Power Reactor Facilities” Effective dose equivalent to the whole body shall not exceed 0.25 Sv outside the site boundary Effective dose equivalent to the thyroid shall not exceed 1.5 Sv for a child outside the site boundary Based on “Examination Guide of Reactor Siting and Guidelines for Interpretation in their Application” Effective dose equivalent to the whole body shall not exceed 0.25 Sv outside the site boundary Effective dose equivalent to the thyroid shall not exceed 3.0 Sv for an adult outside the site boundary The whole population dose shall not exceed 2 3 104 man Sv Based on “Examination Guide of Reactor Siting and Guidelines for Interpretation in their Application”

LWR. Table 2.36 provides the summary of the top-level regulatory criteria for both the HTTR and LWRs in Japan [81]. In addition, according to the International Commission on Radiological Protection (ICRP) recommendation [82], the principle of ALARA (as low as reasonably achievable) was applied to reduce the radiation dose to plant personnel and members of the public around the HTTR as low as reasonably achievable. To meet the Top Level Criteria and apply the principle of ALARA, the basic safety design philosophy of the HTTR was determined based on that of the LWR stipulated in “Guidelines for Safety Design of LWR Power Plant” considering inherent safety characteristics of the HTGR. In order to ensure safety, the well-known fundamental safety functions such as control of reactivity, removal of heat from the core, and confinement of radioactive materials shall be performed in normal and off-normal states. The strategy of the defense in depth [83] that provides a series of level of defense is implemented to ensure that the fundamental safety functions shall be reliably achieved in normal and off-normal states. The

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High Temperature Gas-cooled Reactors

level of the defense in depth consists of prevention of off-normal events, control of off-normal events, and mitigation of off-normal events. The newly considered premise to establish the basic safety design philosophy of the HTTR is that it considers an air ingress accident and following oxidation of the core. The HTTR safety design shall prevent the excessive oxidation of the core and FPs release to the environment.

2.13.3 Safety classification In compliance with the strategy of the defense in depth, all systems having safety functions in the HTTR were classified as Prevention System (PS)-1, PS-2, PS-3, Mitigation System (MS)-1, MS-2, or MS-3 depending on their roles and significance. The definition of each classification is shown as follows: PS-1 Structures, systems, and components whose failure or malfunction has the possibility to cause apparent damage in the core or a large amount of fuel failure. PS-2 Structures, systems, and components whose failure or malfunction does not have the possibility to cause the immediate apparent damage in the core or a large amount of fuel failure, however, has the possibility to cause the excessive amount of FP release outside the site boundary. Structures, systems, and components whose operation is expected in normal operation or anticipated operational occurrences and their malfunctions have the high possibility to damage the core cooling performance. PS-3 Structures, systems, and components are not classified as PS-1 and PS-2, and their failures or malfunctions have the possibility to trigger off-normal events. Every system to keep the amount of FPs in the primary circuit lower than the allowable level MS-1 Structures, systems, and components provided to stop the reactor rapidly, remove the residual heat, prevent the temperature and internal pressure increase of the primary pressure boundary, and prevent the excessive amount of radiation exposure to the members of public outside the site boundary. Other facilities indispensable to safety. MS-2 Structures, systems, and components provided to sufficiently reduce the radiation exposure to the public outside site boundary in the event caused by failure or malfunction of those classified as PS-2. Structures, systems, and components especially significant as countermeasures in offnormal states. MS-3 Structure, systems, and components provided to mitigate the anticipated operational occurrences together with those of classified as MS-1 and MS-2. Everything significant as countermeasures in off-normal states. G

G

G

G

G

G

G

G

G

G

G

Table 2.37 shows the safety classification of the HTTR.

Table 2.37 Safety classification of HTTR [81]. Class

Function

Structure, system, component

PS-1

Primary pressure boundary

Components, pipes in primary circuit except small pipes such as pipes for instrumentation

Prevention of insertion of excessive reactivity

Stand pipe Stand pipe closure

Core constitution

Core structures (fuel block, reflector block, etc.) In-core graphite structure In-core metal structure

Emergency stop of reactor

Control rod system

Maintain subcriticality

Control rod system Reserve shutdown system

Prevention of excessive pressure of primary pressure boundary

Safety valve in primary circuit

Decay heat removal

Auxiliary cooling system Vessel cooling system

Engineered safety system Core cooling

Auxiliary cooling system Vessel cooling system

Containment of FP

Containment vessel Emergency air purification system

Sending signal for engineered safety systems and reactor shutdown system

Engineered safety features actuating system

The other systems having safety-related function

Emergency generator Control room Electric facility Auxiliary plant facility

MS-1

(Continued)

Table 2.37 (Continued) Class

Function

Structure, system, component

PS-2

Containing primary coolant

Primary helium purification system

Storage of radioactive waste

Gaseous radioactive waste treatment system Spent fuel storage pool, cell, rack

Safety handling of fuel

Fuel handling machine

Function related to irradiation test

Irradiation test facility

Closing safety valve at proper pressure

Safety valve in primary circuit

Decrease of FP release

Stack

Postaccident measurement

Postaccident instrumentation

Reactor shut down outside control room

Shut down system outside control room

Containing primary coolant (which is not classified as PS-1 and PS-2)

Instrumentation pipe Primary helium sampling system Primary helium makeup system

Circulation of primary coolant

Primary helium circulator

Storage of radioactive waste

Liquid radioactive waste treatment system Solid radioactive waste treatment system

Cooling of secondary helium circuit during normal operation

Secondary helium cooling system

Maintain differential pressure between primary and secondary helium circuit

Secondary helium makeup system

MS-2

PS-3

(Continued)

Table 2.37 (Continued) Class

MS-3

Function

Structure, system, component

Plant control and instrumentation (except engineered safety features actuating function)

Reactor control system Reactor instrumentation system Process instrumentation system

Plant auxiliary function

Compressed air system for control system

Cooling of vessel cooling system during normal operation

Vessel cooling system

Function related to irradiation test

Irradiation test facility (except ones classified as PS-2)

Prevention of FP release to primary helium coolant

Coated layers of fuel Graphite sleeve for fuel

Purification of primary coolant

Primary helium purification system

Mitigation of reactor power increase

Interlock for control rod withdrawal Interlock for control rod pattern

Mitigation of decreasing coolability in reactor core

Circuit breaker for primary water pump Interlock for inlet temperature of primary water in primary pressurized water cooler Interlock for water flow rate in secondary pressurized water cooler

Mitigation of temperature increase of primary pressure boundary

Frequency converter for helium circulators

Significant function for off-site emergency plan

Sampling system during accident, etc.

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High Temperature Gas-cooled Reactors

2.13.4 Fundamental safety functions unique to HTTR 2.13.4.1 Control of reactivity The reactor is shut down safely and reliably from any operational state using the control rod system. Furthermore, the reserve shutdown system is provided, which is composed of boron-carbide/graphite (B4C/C) pellets. The power control and normal reactor shutdown of the HTTR are achieved with 16 pairs of control rods. The control rod system can achieve subcriticality from any operation state and maintain subcriticality in the cold core conditions even when a pair of control rods sticks at the operational position. In the case of a scram during normal operation, nine pairs of control rods in the replaceable reflector region are inserted at first, and the rest of control rods are inserted after the core is cooled down to prevent exposure of the control rod cladding in a high temperature environment of above 900 C. The core temperature is determined by monitoring the outlet helium gas temperature. A pair of control rods is driven by the one-drive mechanism. The control rods are released from the drive mechanism and inserted by gravity when the reactor is scrammed.

2.13.4.2 Removal of heat from core The main cooling system removes residual heat from the core during a normal reactor shutdown. In addition to the main cooling system, the HTTR has two other residual heat removal systems. The auxiliary cooling system is used for off-normal transients that coolant flow boundary is intact and a VCS is used for accidents that forced circulation of the coolant cannot be maintained. The auxiliary cooling system automatically starts up when the reactor is scrammed in an anticipated operational occurrence and accident in which forced cooling is available, while the main cooling system is stopped. The auxiliary cooling system is classified as a safety system because it has safety function to cool heat transfer tubes in the PPWC in abovementioned off-normal states and keep their temperature lower than the allowable temperature. It consists of two helium gas circulators, the auxiliary water cooler and an affiliated water-cooling system. In terms of the core coolability, the residual heat can be removed by the VCS without using the auxiliary cooling system. However, it is needed from the viewpoint of operational flexibility because it takes a very long time to cool down the core without the auxiliary cooling system. The VCS is used as a residual heat removal system when the forced circulation in the primary cooling circuit is no longer available due to a rupture of the inner pipe or both internal and external pipes in the coaxial double primary pipes. The VCS is also a safety system equipped with two completely independent systems, which are backed up with an emergency power supply. It is operated even during normal operation to cool the reactor shielding concrete wall.

2.13.4.3 Confinement of fission product release The HTTR has multiple barriers to prevent FP release into the environment, fuel coatings, the reactor pressure boundary, the containment vessel, and the reactor building. The ceramic layers surrounding the fuel kernel act as the primary barrier for the FP

Design of High Temperature Engineering Test Reactor (HTTR)

165

release. The integrity of these ceramic layers is sufficiently kept under 1600 C based on several experiments. JAERI carried out the irradiation test and postirradiation tests up to 33,000 MWd/t before the HTTR operation and continues to carry out irradiation tests for the HTTR initial loading fuel up to 70,000 MWd/t. So far, the heating up tests after the irradiation proved that the integrity of the fuel can be sufficiently maintained under 1600 C. It also showed that the fuel failure rate in the range from 1600 C to 1800 C is negligibly small. This is the major specific feature of the HTTR as well as the other HTGRs in the world. While most of the HTGRs constructed or being designed in the world do not have a containment vessel, the HTTR has a containment vessel made of steel. Its functions are to contain FPs and to limit the amount of air ingress into the core. The containment vessel is installed in the reactor building, which acts as the confinement. The confinement is maintained at a slightly negative pressure to the environment by a ventilation and air conditioning system during both normal and off-normal states. The off-site radiation dose limit in such accident as depressurization accident is remarkably reduced by the containment vessel together with the confinement.

2.13.5 Acceptance criteria Acceptance criteria for the HTTR are established fundamentally reflecting the safety requirements for LWR power plants and taking into account major features of HTGRs and the HTTR. Acceptance criteria for the anticipated operational occurrences and the accidents for the HTTR and LWRs are shown in Table 2.38. The maximum fuel temperature is restricted to 1600 C to avoid fuel failure during the anticipated operational occurrences. Criteria for the temperature and the pressure of the primary pressure boundary and the containment vessel are determined considering the following items: (1) materials composing the pressure boundary and the containment vessel shall have stable strength, and their temperature range during the normal operation and abnormal condition is within the temperature range determined by the design code [31], (2) the materials such as 21/4Cr1Mo and Hastelloy XR have sufficient strength below the temperature of 550 C and 1000 C, respectively. However, the margin in their creep rupture strength decreases over 500 C for 21/4Cr1Mo and 980 C for Hastelloy XR. The temperature limits of their materials in anticipated operational occurrences are determined to be 500 C and 980 C, respectively, so that the reactor components can be reused without any repair after an anticipated operational occurrence occurs. In the case of accidents, the core shall not be seriously damaged and shall maintain its geometry for sufficient coolability, that is, (1) the fuels shall be maintained in the graphite fuel block or sleeve and (2) the structural integrity of the graphite support structures such as support posts shall prevent the core from collapsing so as to maintain subcriticality. The radiation exposure is limited to 5 mSv as effective dose equivalent outside the site boundary of the HTTR. For the evaluation of the radiation exposure, external gamma ray exposure from the radioactive cloud containing noble gases and iodine, internal exposure by inhalation from the radioactive cloud, direct external gamma ray exposure, and external skyshine gamma ray exposure from FPs such as Cesium contained in the containment vessel are considered. The total radiation exposure, which is the sum of these exposures, shall be lower

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High Temperature Gas-cooled Reactors

Table 2.38 Acceptance criteria for HTTR [81]. HTTR

LWR

(1) Anticipated operational occurrence G

The peak fuel temperature shall be less than 1600 C

G

Pressure on reactor pressure boundary is less than 1.1 times of maximum pressure in service

G

Maximum temperature of reactor pressure boundary

G

Minimum critical heat flux (MCHF) or MCHF ratio shall not exceed the limited value

G

Fuel cladding shall not fail mechanically

G

Fuel enthalpy shall not exceed the limited value

G

Pressure on reactor pressure boundary is less than 1.1 times of maximum pressure in service

G

Minimum critical heat flux (MCHF) or MCHF ratio shall not exceed the limited value

G

Fuel cladding shall not fail mechanically

G

Fuel enthalpy shall not exceed the limited value

G

Pressure on reactor pressure boundary is less than 1.2 times of maximum pressure in service

21/4Cr1Mo steel ,500 C Austenite stainless steel ,600 C Hastelloy XR ,980 C (2) Accident G

The reactor core shall not be seriously damaged and can be cooled sufficiently

G

The reactor core shall not be seriously damaged and can be cooled sufficiently

G

Maximum temperature of reactor pressure boundary

21/4Cr1Mo steel ,550 C Austenite stainless steel ,650 C Hastelloy XR ,1000 C G

Maximum pressure on containment boundary is less than maximum pressure in service

G

Maximum pressure on containment boundary is less than maximum pressure in service

G

No significant risk of radiation exposure to public

G

No significant risk of radiation exposure to public

Design of High Temperature Engineering Test Reactor (HTTR)

167

than the limit of 5 mSv. Based on ICRP, in special case, 5 mSv of annual radiation exposure for the public is acceptable, though a limit of 1 mSv is recommended. For accidents, which have small frequencies of those occurrences, the value of 5 mSv is applied to judge the significant risk of radiation exposure for the public. The nearest site boundary from the HTTR facility is about 200 m and the site is about 5 km far from the center of Oarai town having about 20,000 inhabitants.

2.13.6 Selection of events Abnormal events to be postulated as anticipated operational occurrences and accidents have been selected considering their frequencies of occurrence and based on the investigation of main causes, which affect each item of the acceptance criteria identified for the HTTR, that is, (1) fuel temperature, (2) core damage, (3) temperature of reactor coolant pressure boundary, (4) pressure at reactor coolant pressure boundary, (5) pressure at containment vessel boundary, and (6) risk of radiation exposure for the public. The initiating abnormal events have been classified into similar event groups according to “Examination Guide for Safety Evaluation of Light Water Nuclear Power Reactor Facilities.” Then the most severe events with respect to the acceptance criteria within each similar event group are selected as the representative postulated events. Examples of the selection of anticipated operational occurrences and accidents are shown in Figs. 2.91 and 2.92. The main causes affecting the fuel temperature are increase of power and decrease of heat removal in the core. The increase of power is caused by reactivity addition, in which four event groups are postulated, namely, (1) malfunction of the reactivity control system, (2) malfunction of the experimental facility, (3) increase of the primary coolant flow, and (4) increase of heat removal by the secondary cooling system. Two initiating events are considered as the malfunction of the reactivity control system, namely (1) abnormal control rod withdrawal and (2) abnormal control rod insertion. The abnormal control rod withdrawal is the more severe event with respect to the fuel temperature, and is selected as the representative postulated event. The representative postulated events concerning other acceptance criteria are selected in the same way. The postulated events considered in the safety evaluation of the HTTR as anticipated operational occurrences and accidents are listed in Table 2.39. A major accident and a hypothetical accident are evaluated to ensure the safety of the public in the case of serious accidents. “Major accidents” are postulated assuming the occurrence of the worst-case accident from a technical standpoint considering the reactor characteristics and engineered safety features. “Hypothetical accidents” are postulated assuming the occurrence of an accident more serious than a “major accident,” which is unlikely to occur from a technical standpoint, and shall be based on the assumption that one or more engineered safety features fail to function. The acceptance criteria are established in “Examination Guide of Reactor Siting and Guidelines for Interpretation in their Application” as follows: (1) effective dose equivalent to whole body shall not exceed 0.25 Sv in a major accident or a hypothetical accident; (2) effective dose equivalent to thyroid shall not exceed 1.5 Sv for a child in a major accident and 3.0 Sv for an adult in a hypothetical accident; (3) whole population dose shall not exceed 2 3 104 man Sv in a hypothetical accident. A double-ended rupture of coaxial double pipes of the primary cooling system

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High Temperature Gas-cooled Reactors

Incrrease of power

Malfunction of reactivity control system

Abnormal control rod withdrawal Abnormal control rod insertio

Malfunction of experimental facility

Abnormality of irradiation specimeen

Increase of primary coolant flow

Abnorm mal rotaion increase of primary pressurized watercirculator, or interm mediate heat exchangeer circulator, Abnor mal start of auxiliary c ooling system, etc.

Increase of heat removal by secondary cooling system

Abnorm mal rotaion increase of secondary pressuri zed water cooler circulator, etc.

Decrease of primary coolant flow

Coast doown of primary pressuurized water cooler or interm mediate heat exchanger circulator, etc. Abnorm al control rod insertion

Decrease of heat removal by pressurized water cooling system

Stop of air cooler, abnormal close of isolation valve, etc.

Decrease of heat removal by secondary helium cooling system

Coast down of secondary helium ccirculator, etc.

Reactivity addition

Incrrease of fuel tem mperature

Increase of core c coolant flow

Incrrease of heat rem moval in core

Increase of core c coolant temperature

Figure 2.91 Example of selection process of anticipated operational occurrence [81]. Increase of power

Reactivity addition

Malfunction of reactivity control system

Rupture of standpipe

Decrease of heat removal from core

Local decrease of core coolant flow

Decrease of coolant flow in fuel channel

Channel blockage in fuel block

Decrease of core coolant flow

Rupture of inner pipe of co-axial double pipes in primary cooling system, Rupture of inner pipe of co-axial double pipes in auxiliary cooling system, etc.

Loss of primary coolant

Rupture of pipes of primary cooling system, Rupture of pipes of auxiliary cooling system, etc.

Decrease of secondary helium

Rupture of inner pipe of co-axial double pipes in secondary cooling system, etc.

Loss of secondary helium flow

Rupture of co-axial double pipes in secondary cooling system, etc.

Decrease of heat removal by pressureized water cooler

Rupture of pipes in pressurized water cooler, etc.

Air ingress

Rupture of co-axial double pipes in primary cooling system in auxiliary cooling system, etc.

Water ingress

Rupture of heat tube in primary water cooler, in auxiliary cooling system, etc.

Increase of fuel temperature

Core damage

Increase of core inlet temperature

Loss of core support function Graphite corrosion Drop of fuel element

Figure 2.92 Example of selection process of accident [81].

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Table 2.39 Selected events [81]. (1) Anticipated operational occurrence Abnormal control rod withdrawal under subcritical condition Abnormal control rod withdrawal during rated operation Decrease in primary coolant flow rate Increase in primary coolant flow rate Decrease in heat removal by secondary cooling system Increase in heat removal by secondary cooling system Loss of off-site electric power Abnormality of irradiation specimens and experimental equipment Abnormality during safety demonstration tests (2) Accident Channel blockage in fuel block Rupture of inner pipe of concentric pipe in primary cooling system Rupture of inner pipe of concentric pipe in secondary cooling system Rupture of concentric pipe in secondary cooling system Rupture of pipe in pressurized water-cooling system Rupture of concentric pipe in primary cooling system Rupture of heat tube in pressurized water cooler Rupture of pipe in primary coolant purification system Rupture of pipe in processing facilities of radioactive gaseous waste Rupture of sweep gas pipe in irradiation test facilities Rupture of stand pipe

(depressurization accident) is postulated for the HTTR as the major and hypothetical accidents with respect to the risk of radiation exposure for the public.

2.13.7 Safety evaluation technologies This section summarizes “Verified analysis codes accepted to licensing” [84]. Analytical tools for the safety evaluation were developed. All of them were validated by comparison between experimental results and analytical ones for the safety evaluation of the HTTR. Table 2.40 shows the representative events in the safety evaluation of the HTTR and analytical tools. Analytical code is as follows:

Table 2.40 Selected events and analytical codes. Event name

BLOOSTJ2

THYDEHTGR

TACNC

RATSAM6

COMPAREMOD1

GRACE

OXIDE3F

FLOWNET/ TRUMP

Anticipated operational occurrence Abnormal control rod withdraw under subcritical condition

Y

Abnormal control rod withdraw under rated operation

Y

Decrease in primary coolant flow rate

Y

Increase in primary coolant flow rate

Y

Decrease in heat removal by secondary cooling system

Y

Increase in heat removal by secondary cooling system

Y

Loss of off-site electric power Abnormality of irradiation specimens and experimental equipment Abnormality during safety demonstration tests

Y Y Y

(Continued)

Table 2.40 (Continued) Event name

BLOOSTJ2

THYDEHTGR

TACNC

RATSAM6

COMPAREMOD1

GRACE

OXIDE3F

FLOWNET/ TRUMP

Accident Channel blockage in fuel block

Y

Rupture of inner pipe of coaxial double pipes in primary cooling system

Y

Rupture of inner pipe of coaxial double pipes in secondary cooling system

Y

Rupture of coaxial double pipes in secondary cooling system

Y

Rupture of pipe in pressurized water-cooling system

Y

Rupture of coaxial double pipes in primary cooling system

Y

Rupture of heat tube in pressurized watercooling system

Y

Y

Y

Y

Rupture of pipe in primary coolant purification system

Y

Y

Rupture of pipe in processing facilities of radioactive gaseous waste Rupture of sweep gas pipe in irradiation test facilities Rupture of stand pipe

Y Y

Y

Y

Y

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High Temperature Gas-cooled Reactors

The BLOOST-J2 code [85] is used to analyze the effects of reactivity and flow rate change on the reactor power and temperatures of the core. The BLOOST-J2 code was modified from BLOOST5 code [86] so as to adopt the configuration of the HTTR. The validation of the BLOOST-J2 code was conducted by comparing analytical results with the data of control rod withdrawal/insertion experiments with Fort St. Vrain (FSV) at 50% of rated power. The THYDE-HTGR code [87] is used to analyze plant dynamics of the HTTR. The THYDE-HTGR code was modified from THYDE [88] code to treat helium gas behavior in transient conditions. The THYDE code was validated by a comparison with various experiments such as LOFT experiments [89]. A new function to evaluate thermal and hydraulic transient of helium gas was added in the THYDE-HTGR code. The other functions were the same as that of the THYDE code. The validation of the THYDE-HTGR code was conducted by comparing analytical results with the data of control rod withdrawal/insertion experiments with FSV at 50% of rated power. Thermal and hydraulics behavior of helium gas was validated by comparing experimental results obtained in Engineering Research Association of Nuclear Steelmaking (ERANS) and analytical ones. The TAC-NC code [90] modified from TAC-2D [91] is used to calculate transient thermal and hydraulic characteristics in the core during LOFC accident such as the depressurization accident. The TAC-2D code was used for various calculations, and heat transfer calculations by conduction, convection, and radiation have a sufficient reliability. The function to analyze heat transfer by natural circulation in the core is added in the TAC-NC code. This function was validated by comparing with an air ingress experiment, which simulated a rupture of the coaxial double primary pipes of the HTTR. The RATSAM6 code [92] is used to calculate the amount of mass and energy released from the reactor into the containment vessel with consideration given to heat transfer during the rupture of the coaxial double primary pipes. The validation of the code was performed by comparing the analytical results with experimental results, which are obtained by using a 1/8 scaled apparatus simulating the primary cooling system of the Colder Hall-type Reactor. The COMPARE-MOD1 code [93] is used to calculate pressure and temperature behavior in each compartment of the containment vessel during the depressurization accident. The code was certified by the US Nuclear Regulatory Committee as a code for safety analysis to calculate pressure and temperature behavior in the containment vessel. The GRACE code [94] is used to calculate axial and radial distributions of oxidation of graphite materials and concentration distribution of oxygen in a mixed gas of air and helium by analyzing the oxidation reaction between ingressed air and the graphite structures. The validation of the code was performed by using results of a graphite oxidation experiment. Input conditions for the code such as mass transfer coefficient were obtained from the heat transfer correlations obtained in the experiments. The OXIDE-3F code [95] is used to analyze the oxidation reaction of the graphite materials with steam ingress in the core by a rupture of the heat transfer tubes of the PPWC. The method to calculate the rate of graphite oxidation in the OXIDE-3F code is basically the same as that of the GRACE code.

Design of High Temperature Engineering Test Reactor (HTTR)

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The FLOWNET/TRUMP code [96] is used to calculate the temperature distribution in the fuel block when a coolant channel was blocked. The code is the combination of the FLOWNET and TRUMP codes. The FLOWNET code is a one-dimensional flow network evaluation code while the TRUMP code is a three-dimensional heat conduction code. The validation of the code was performed by comparing results of the uniform and nonuniform power distribution tests, which were carried out using the multichannel test rig of the HENDEL.

2.13.8 New safety criteria After the 2011 off the Pacific coast of Tohoku Earthquake, Nuclear Regulatory Authority established a more stringent new regulation standard. On the new regulatory standards, the safety designs for the strengthened natural phenomena such as earthquakes and volcanoes, fire, and internal flooding were required. Furthermore, the mitigation measures against the beyond design basis accidents were required. The HTTR is considering changing the safety classification by considering safety functions such as radiant core cooling even if a loss of a forced cooling accident occurs. The safety review for the HTTR to confirm the conformity to the new regulation standard is undergoing.

References [1] S. Shiozawa, et al., Overview of HTTR design features, Nucl. Eng. Des. 233 (2004) 1121. [2] Y. Tachibana, et al., Procedure to prevent temperature rise of primary upper shielding in high temperature engineering test reactor, Nucl. Eng. Des. 201 (2000) 227238. [3] K. Takamatsu, et al., High-temperature continuous operation of the HTTR, Trans. At. Energy Soc. Jpn. 10 (2011) 290300. [4] M. Goto, et al., Long-term high-temperature operation in the HTTR (2) Core physics, in: Proceedings of the HTR 2010, Prague, Czech Republic, 2010. [5] S. Ueta, et al., Development of high temperature gas-cooled reactor (HTGR) fuel in Japan, Prog. Nucl. Energy 53 (2011) 788793. [6] S. Hamamoto, et al., Chemical characteristics of helium coolant of HTTR (High Temperature engineering Test Reactor), in: Proceedings of the HTR 2012, Tokyo, Japan, 2012. [7] A. Shimizu, et al., Development of operation and maintenance technology of HTTR (High Temperature engineering Test Reactor), in: Proceedings of the HTR 2012, Tokyo, Japan, 2012. [8] Y. Tachibana, et al., Test Plan Using the HTTR for Commercialization of GTHTR300C, Japan Atomic Energy Agency, JAEA-Technology 2009-063, 2009. [9] K. Takamatsu, et al., Experiments and validation analyses of HTTR on loss of forced cooling under 30% Reactor power, J. Nucl. Sci. Technol. 51 (2014) 14271443. [10] M. Ono, et al., Comprehensive seismic evaluation of HTTR against the 2011 Off the Pacific Coast of Tohoku Earthquake, ASME J. Nucl. Rad. Sci. 4 (2) (2018). NERS-161104. [11] T. Nishihara, et al., Excellent Features of Japanese HTGR Technologies, Japan Atomic Energy Agency, JAEA-Technology 2018-004, 2018.

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[35] F. Garofalo, et al., Straintime, ratestress, and ratetemperature relations during large deformations in creep, in: Proceedings of the Joint International Conference on Creep, Institute Mechanical Engineering, London, 1963, pp. 131. [36] M.K. Booker, et al., Mechanical property correlations for 21/4Cr1Mo steel in support of nuclear reactor systems design, Int. J. Press. Vessel. Pip. 5 (1977) 181204. [37] K. Hada, Influence of variations in creep curve on creep behavior of a hightemperature structure, Nucl. Eng. Des. 97 (1986) 279296. [38] T. Iyoku, et al., Design of core components, Nucl. Eng. Des. 233 (2004) 7179. [39] J. Sumita, et al., Reactor internals design, Nucl. Eng. Des. 233 (2004) 8188. [40] K. Sawa, et al., Research and development on HTGR fuel in the HTTR project, Nucl. Eng. Des. 233 (2004) 163172. [41] K. Hayashi, et al., Design Criteria, Production and Total Integrity Assessment of Fuels of the High Temperature Engineering Test Reactor, Japan Atomic Energy Research Institute, JAERI-M 89161, 1989. [42] K. Fukuda, et al., Research and Development of HTGR Fuel, Japan Atomic Energy Research Institute, JEARI-M 89-007, 1989. [43] K. Sawa, et al., Fabrication of the first-loading fuel of the high temperature engineering test reactor, J. Nucl. Sci. Technol. 36 (1999) 683690. [44] K. Minato, et al., Fission product palladium-silicon carbide interaction in HTGR fuel particles, J. Nucl. Mater. 172 (1990) 184196. [45] K. Hayashi, et al., Assessment of Fuel Integrity of the High Temperature Engineering Test Reactor (HTTR) and Its Permissible Design Limit, Japan Atomic Energy Research Institute, JAERI-M 89162, 1989. [46] M. Ishihara, T. Iyoku, Development of graphite design philosophy, Nucl. Eng. Des. 233 (2004) 251260. [47] T. Iyoku, et al., Development of thermal/irradiation stress analytical code VIENUS for HTTR graphite block, J. Nucl. Sci. Technol. 28 (10) (1991) 921931. [48] M. Ishihara, et al., Development of irradiation-induced stress analysis code system for graphite components in gas-cooled reactor, in: Proceedings of the 12th International Conference on Structural Mechanics in Reactor Technology (SMiRT 12), Stuttgart, Germany, 1993, C08/1. [49] T. Iyoku, et al., Seismic response of the high-temperature engineering test reactor core bottom structure, Nucl. Technol. 99 (1992) 169176. [50] Nuclear Safety Commission, Guidelines for Aseismic Design of Nuclear Power Plants, 1981. [51] K. Iigaki, et al., Seismic design, Nucl. Eng. Des. 233 (2004) 5970. [52] T. Iyoku, et al., R&D on core seismic design, Nucl. Eng. Des. 233 (2004) 225234. [53] T. Iyoku, et al., Seismic study of High-Temperature Engineering Test Reactor core graphite structures, Nucl. Technol. 99 (1992) 158168. [54] T. Ikushima, SONATINA-2V: A Computer Program for Seismic Analysis of the TwoDimensional Vertical Slice HTGR Core, Japan Atomic Energy Research Institute, JAERI-1279, 1982. [55] M. Ishihara, T. Iyoku, M. Futakawa, Evaluation of aseismic integrity in HTTR core-bottom structure. III. Structural integrity of core support post component, Nucl. Eng. Des. 148 (1994) 91100. [56] M. Ishihara, T. Iyoku, M. Futakawa, Evaluation of aseismic integrity in HTTR corebottom structure. IV. Structural integrity of connecting elements between graphite components, Nucl. Eng. Des. 158 (1995) 8395. [57] T. Furusawa, et al., Cooling system design and structural integrity evaluation, Nucl. Eng. Des. 233 (2004) 113124.

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[58] M. Shindo, T. Kondo, Studies on improving compatibility of nickel-base alloys with high temperature helium-cooled reactor (VHTR) environment, in: BNES Conference, Gas-Cooled Reactors Today, Bristol, 1982. [59] T. Nishihara, et al., Demonstration test of hydrogen production using high temperature gas-cooled reactor, in: Proceedings of the 15th World Hydrogen Energy Conference, Yokohama, Japan, 2004. [60] Y. Tachibana, et al., Reactivity control system of the high temperature engineering test reactor, Nucl. Eng. Des. 233 (2004) 89101. [61] S. Maruyama, et al., Temperature Analysis of Control Rod for HTTR, Japan Atomic Energy Research Institute, JAERI-M 90-104, 1990. [62] I. Nishiguchi, et al., General Criteria for the Structural Design of the HTTR Control Rods, Japan Atomic Energy Research Institute, JAERI-M 90-152, 1990. [63] K. Watanabe, T. Kondo, Y. Ogawa, Postirradiation tensile and creep properties of heatresistant alloys, Nucl. Technol. 66 (1984) 630638. [64] National Research Institute for Metals (NRIM) Creep Data Sheet No. 26 A, Data Sheets on the Elevated-Temperature Properties of Iron Base 21Cr32NiTiAl Alloy Plates for Corrosion and Heat Resistant Applications (NCF 800H-P), National Research Institute for Metals, Tokyo, 1983. [65] National Research Institute for Metals (NRIM) Creep Data Sheet No. 27 A, Data Sheets on the Elevated-Temperature Properties of Iron Base 21Cr32NiTiAl Alloy Tubes for Heat Exchanger Seamless Tubes (NCF 800H-TB), National Research Institute for Metals, Tokyo, 1983. [66] Y. Monma, et al., Assessment of elevated-temperature property data for alloy 800H, Trans. Natl. Res. Inst. Met. 26 (3) (1984) 3347. [67] J. Fukakura, F. Matsumoto, T. Araki, Effect of strain hold time on fatigue life of alloy 800H at high temperatures, Trans. Jpn. Soc. Mech. Eng. 57 (540) (1991) 17001705. [68] R. Hino, Y. Miyamoto, H. Fukushima, Reliability test on control rod driving mechanism of HTTR with HENDEL, J. At. Energy Soc. Jpn. 33 (1991) 685694. [69] A.L. Edwards, TRUMP: A Computer Program for Transient and Steady State Temperature Distribution in Multidimensional Systems, Lawrence Livermore Laboratory report, UCRL14754, rev. 3, 1972. [70] M. Hirano, K. Hada, Development of THYDE-HTGR: Computer Code for Transient Thermal-Hydraulics of High Temperature Gas-Cooled Reactor, Japan Atomic Energy Research Institute, JAERI-M 90-071, 1990. [71] Y. Tachibana, et al., Integrity assessment of the High Temperature Engineering Test Reactor (HTTR) control rod at very high Temperatures, Nucl. Eng. Des. 172 (1997) 93102. [72] Y. Shimakawa, et al., The plant dynamics analysis code ASURA for the high temperature engineering test reactor (HTTR), Specialist’s meeting on uncertainties in physics calculations for gas cooled reactor cores, Villigen, Switzerland, 1990. [73] K. Saito, et al., Instrumentation and control system design, Nucl. Eng. Des. 233 (2004) 125133. [74] N. Sakaba, et al., Leak-tightness characteristics concerning the containment structures of the HTTR, Nucl. Eng. Des. 233 (2004) 135145. [75] M. Kondo, et al., Leakage Rate Test for Reactor Containment Vessel of HTTR, Japan Atomic Energy Agency, JAEA-Testing 2006-002, 2006. [76] K. Iigaki, et al., Performance tests of reactor containment structures of the HTTR, in: Proceedings of the 16th International Conference on Structural Mechanics in Reactor Technology (SMiRT 16), Washington, DC, 2001.

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[77] T. Aono, et al., Maintenance and Management of Emergency Air Purification System in HTTR, Japan Atomic Energy Agency, JAEA-Testing 2006-004, 2006. [78] N. Sakaba, et al., Short design descriptions of other systems of the HTTR, Nucl. Eng. Des. 233 (2004) 147154. [79] Y. Inagaki, et al., Cooling performance of helium-gas/water coolers in HENDEL, Nucl. Eng. Des. 146 (1994) 301309. [80] S. Nakagawa, et al., Development of the Unattended Spent Fuel Flow Monitoring Safeguards System (UFFM) for the High Temperature Engineering Test Reactor (HTTR) (Joint research), Japan Atomic Energy Agency, JAEA-Technology 2007-003, 2007. [81] K. Kunitomi, et al., Safety design, Nucl. Eng. Des. 233 (2004) 4558. [82] ICRP, Recommendations of the International Commission on Radiological Protection, ICRP Publication 26, Program Press, Elmsford, NY, 1977. [83] International Atomic Energy Agency (IAEA), Basic Safety principles for Nuclear Power Plants 75-INSAG-3 Rev. 1, INSAG-12, 1999. [84] K. Kunitomi, et al., Safety evaluation of the HTTR, Nucl. Eng. Des. 233 (2004) 235249. [85] S. Nakagawa, et al., Core Dynamics Analysis Code for High Temperature Gas Reactor BLOOST-J2, JAERI-M 89-0132, 1989. [86] M. Merrill, BLOOST-5: A Combined Reactor Kinetics—Heat Transfer Code for the IBM-7044; Preliminary Description, General Atomics, GAMD-6644, 1965. [87] M. Hirano, et al., Development of THYDE-HTGR: Computer Code for Transient Thermal-Hydraulics of High-Temperature Gas-Cooled Reactor, Japan Atomic Energy Research Institute, JAERI-M 90-071, 1990. [88] Y. Asahi, et al., THYDE-P2 Code: RCS (Reactor-Coolant System) Analysis Code, Japan Atomic Energy Research Institute, JAERI 1300, 1986. [89] M. Hirano, Analysis of LOFT Small Break Experiment L3-1 with THYDE-P Code: CSNI International Standard Problem No. 9 and THYDE-P Sample Calculation Run 50, JAERI-M 82-008, 1982. [90] K. Kunitomi, et al., Two-Dimensional Thermal Analysis Code “TAC-NC” for High Temperature Engineering Test Reactor and Its Verification, Japan Atomic Energy Research Institute, JAERI-M 89-001, 1989. [91] S.S. Clark, J.F. Petersen, TAC2D: A General Purpose Two-Dimensional Heat Transfer Computer Code—Mathematical Formulations and Programmer’s Guide, General Atomics, GA-9262, 1969. [92] R.K. Deremer, et al., RATSAM: A Computer Program to Analyze the Transient Behavior of the HTGR Primary Coolant System During Accidents, General Atomics, GA-A-13705, 1977. [93] R.G. Gido, et al., COMPARE-MOD1: A Code for the Transient Analysis of Volumes with Heat Sinks, Flowing Vents and Doors, Los Alamos, LA-7199-MS, 1978. [94] H. Kawakami, Air oxidation behavior of carbon and graphite materials for HTGR, Tanso 124 (1986) 2633. [95] M.B. Peroominan, OXIDE-3: A Computer Code for Analysis of HTGR Steam or Air Ingress Accidents, General Atomics, GA-A 12493, 1975. [96] S. Maruyama, et al., Verification of Combined Thermal-Hydraulics and Heat Conduction Analysis Code FLOWNET/TRUMP, Japan Atomic Energy Research Institute, JAERI-M 88-173, 1988.

R&D on components

3

Jun Aihara1, Minoru Goto1, Yoshiyuki Inagaki1, Tatsuo Iyoku1, Kazuhiko Kunitomi1, Tetsuo Nishihara1, Nariaki Sakaba1, Taiju Shibata1, Junya Sumita1, Yukio Tachibana1, Shoji Takada1, Tetsuaki Takeda2 and Shohei Ueta1 1 Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan, 2Graduate Faculty of Interdisciplinary Research, Research Faculty of Engineering, Department of Mechanical Engineering, University of Yamanashi, Yamanashi, Japan

The following R&D on main components were carried out for the design, safety review, and performance evaluation of the HTTR: 1. Fuel The design criteria and specifications for the HTTR, and fabrication process are described. Before loading to the HTTR, the fuel performance against high temperature, irradiation, etc. was investigated by the Oarai Gas Loop-1 and capsule irradiation tests, which were installed at the Japan Materials Test Reactor (JMTR). 2. Core components and reactor internals Thermal hydraulic tests on the core and core bottom structure (CBS) of the HTTR were carried out with the helium engineering demonstration loop (HENDEL) under simulated reactor operating conditions. In the test section simulating a fuel stack of the core, thermal and hydraulic performances of helium gas flowing through a fuel block were investigated for thermal design of the HTTR core. In the test section simulating the CBS, demonstration tests were performed to verify the structural integrity of graphite and metal components, seal performance against helium gas leakage among the graphite permanent blocks and thermal mixing performance of helium gas. 3. Passive cooling system Experimental and numerical studies were carried out to investigate the effects of the natural convection of superheated gas as well as those of the standpipes on the temperature distributions of the components and the heat removal performance in the water-cooling panel system for decay heat removal, and to verify reliability of the design and evaluation methods. 4. Intermediate heat exchanger Experimental and numerical studies with partial models were carried out about the following items: creep collapse of the tube against external pressure, creep fatigue of the tube against thermal stress, seismic behavior of the tube bundles, thermal hydraulic behavior of the tube bundles, and in-service inspection technology of the tube. 5. Basic feature of air ingress during primary pipe rupture accident Experimental and numerical studies were performed on the combined phenomena of molecular diffusion and natural circulation in a two-component gas system (N2-He) in a reverse U-shaped

High Temperature Gas-cooled Reactors. DOI: https://doi.org/10.1016/B978-0-12-821031-4.00003-8 © 2021 Elsevier Inc. All rights reserved.

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tube simulating coolant passages in a reactor. The behavior of air ingress was clarified and the numerical method to predict the behavior was developed.

3.1

Fuel

3.1.1 Introduction In the HTGRs, two main fuel element concepts are presently in use, the spherical fuel element and the block-type fuel element. In both concepts, the high temperature heat supply and inherent safety features of HTGRs are mainly achieved using refractory TRISO (tri-structural isotropic)-coated fuel particles (CFPs), where the fuel microsphere (kernel) is coated with the low-density carbon buffer, the inner isotropic high-density carbon (IPyC), the silicon carbide (SiC), and the outer isotropic high-density carbon (OPyC) layers in this order from within. The HTTR applies the TRISO-CFP with uranium dioxide (UO2) kernel. This is dispersed in the graphite matrix and sintered to form a fuel compact as shown in Fig. 3.1. Fuel rods are inserted into vertical holes bored in the graphite block. Then configuration of the fuel rod of each HTTR and GTHTR300 is quite different; the fuel compact of the HTTR is contained in a graphite sleeve to form a fuel rod. Table 3.1 summarizes major specifications of the HTTR fuel [2]. In the field of HTGR fuel, JAEA has carried out a lot of research and development works in the frame of the HTTR Project. Since 1960s, fuel fabrication technologies were developed with the collaboration of the Nuclear Fuel Industries, Ltd. (NFI), and the first-loading fuel of the HTTR was successfully fabricated in December 1997. Fuel performance under normal operation and accident transient conditions was also investigated by the Oarai Gas Loop-1 and capsule irradiation tests, which were installed at the JMTR, in the ranges of temperature and burnup required for the HTTR. The fuel performance and fission product behavior are under investigation through the HTTR operation. For upgrading of HTGR technologies, JAEA has also developed an extended burnup TRISO-CFP, and an advanced type of CFP, where ZrC replaces SiC, in order to raise the operating temperature.

Fuel kernel /UO2 1st buffer layer/Low-dense PyC 3rd layer/SiC

2nd & 4th layers/High dense PyC Coated fuel particle

Figure 3.1 HTTR fuel [1].

Fuel compact

Fuel rod

Fuel block

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181

Table 3.1 Major specifications of HTTR fuels [2]. Fuel kernel

Fuel compact

Diameter (μm)

600 6 55

Density (g/cm3) 235 U enrichment (%)

10.63 6 0.26 3.49.9

Impurity (ppm EBCa) Coating layers Buffer layer thickness (μm) IPyC layer thickness (μm) SiC layer thickness (μm) OPyC layer thickness (μm)

#3 60 6 12 30 6 6 2511 0 45 6 6

Buffer layer density (g/cm3)

1.10 6 0.10

PyC layers density (g/cm3) SiC layer density (g/cm3) OPTAFb of PyC layers

1:8510:10 20:05 $ 3.20 # 1.03

CFPs packing fraction (vol.%) Impurity (ppm EBCa) Exposed uranium fraction SiC-failure fraction Outer diameter (mm) Inner diameter (mm) Height (mm) Matrix density (g/cm3) Compressive strength (N)

30 6 3 #5 # 1.5 3 1024 # 1.5 3 1023 26.0 6 0.1 10.0 6 0.1 39.0 6 0.5 1.70 6 0.05 $ 4900

Fuel rod Uranium content (gU) Total length (mm) Fuel compact stack length (mm)

188.58 6 5.6 577 6 0.5 $ 544

CFP Diameter (μm) Sphericity a

920150 230 # 1.2

Equivalent boron content. Optical anisotropy factor.

b

3.1.2 Related research and development for fuel design In the safety design of the HTGR fuels, it is important to retain fission products within particles so that their release to the primary coolant does not exceed an acceptable level. From this point of view, the basic design criteria for the fuel are to minimize the failure fraction of as-fabricated fuel coating layers and to prevent significant additional fuel failures during the operation [2].

3.1.2.1 Limitation for as-fabricated fuel failure fraction Small fractions of the particles with defective coating layers are present during the fabrication process. Among several modes of defective coating layers, a defective SiC coating layer is the most harmful from the standpoint of fission product retention. In the fabrication of the first-loading fuel of the HTTR, as-fabricated failure fraction was limited less than 0.2% in terms of the sum of exposed uranium and SiC defects. The value of 0.2% was determined from the viewpoint of limit of off-site exposure during normal operation considering fission gas transport from primary coolant to the

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environment [3]. The safety criteria for the HTTR fuel were settled as described in the paper [4]. Corresponding to the criteria, the following research and develop works were carried out to confirm safety characteristics of the fuel [5].

3.1.2.2 Kernel migration Temperature gradients in CFPs due to extreme operating conditions lead to carbon transport from the hot side to the cold side of the IPyC layer causing a migration of the fuel kernel toward the hot side, which is called as the amoeba effect. The distance of kernel migration shall not exceed the thickness of the first buffer layer plus the second IPyC layer to avoid failure of the SiC layer. The experiments on the amoeba effect were carried out by the capsule irradiation where a temperature gradient of about 150 C/cm was imposed on the CFPs embedded in graphite disks. The CFPs were irradiated at a temperature of 1600 C and 1700 C [6]. The kernel migration in the CFPs was measured by X-ray radiography in the post irradiation examination. Experimental data on kernel migration distances from irradiation tests were gathered as shown in Fig. 3.2 [7]. A comparison of the design and the experimental values revealed that failure of the CFPs caused by the amoeba effect did not occur during the fuel-life time since a maximum range of the kernel migration was estimated to be less than 55 μm.

3.1.2.3 Palladium (Pd)SiC interaction The interaction of palladium with SiC layer was experimentally investigated by measuring the depth of penetration of the resulting intermetallic compound into the SiC layer. The penetration depth of the Pd/SiC interaction should not exceed the thickness Temperature (°C) 1900 1800 1700 1600 1500

1400

1300

1200

1100

10–4

Kernel migration rate, KMR (μm/s)

Design equation KMR=2 ×10 −4 exp −

10–5

14,800 1 dT T T 2 dr

10–6

10–7

Sample : 4%eU 10–8 4.5

5.0

5.5

6.0 6.5 104/T (K–1)

7.0

7.5

Figure 3.2 Experimental data on kernel migration distances from irradiation tests [2].

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183

of the SiC layer because the fully penetrating Pd/SiC interaction is thought to lead to the loss of fission product retention in the SiC coating layer. The PdSiC interaction had been investigated intensively in post irradiation experiment where penetration depth of the intermetallic compound was observed in the sectioned SiC coating layer of the irradiated CFPs by electron-probe microanalyzer (EPMA) and ceramography. Through the experiments, a rate-limiting step in this attack was considered to be supplied of Pd from the kernel, since the PdSiC interaction and diffusion of Pd through the buffer and inner PyC layer were fast compared with the release from the kernel. The amount of released Pd was calculated with relations of Pd diffusion in UO2 kernels, irradiation time, temperature, and burnup. Fig. 3.3 shows the relation between the observed PdSiC reaction depth and the calculated palladium amount released from the kernel obtained from the experiments [2,7]. The maximum reaction depth depends on the calculated amount of palladium released from the kernels. It was found that a relationship between the maximum penetration depth and the amount of released Pd was expressed by the cubic root function. As shown in the figure, the maximum penetration depth of the HTTR fuel during the fuel-life time was evaluated to be about 11 μm, which was less than a half of the SiC layer thickness (25 μm), therefore proving integrity of the CFPs.

3.1.2.4 Confirmation up to maximum burnup To satisfy the nuclear design requirement, the maximum value of the average burnup in a fuel element must not be over 33 GWd/t [2]. The performance of the HTTR

Pd–SiC interaction depth (μm)

12

Design curve used for safety analysis of the HTTR fuel

10 8 6 4 2 0

0

0.5 1.0 1.5 2.0 Calculated amount of Pd released from kernel (1015 atoms)

Figure 3.3 Penetration depth is described against the amount of Pd released from the kernel [2].

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High Temperature Gas-cooled Reactors

2000 Peak temp. at trandient test Peak temp. in irrad. test at steady-state power Average temp. during whole irrad. period

Irradiation temperature (°C)

1800

1600

Maximum condition in the HTTR

1400

1200

1000

0

10

20 30 Burnup (GWd/t)

40

50

Figure 3.4 Range of fuel irradiation tests [2].

fuel under normal operating conditions up to the maximum burnup in HTTR design was tested in both the Oarai Gas Loop-1 and capsule irradiation experiments. In situ fission gas release from the fuel was deeply related to the irradiation performance of the fuel. To evaluate the irradiation performance, fractional releases to birth ratio (R/B) of short-lived fission gases from the fuel were determined by measuring the fission gas concentration in the primary helium gas loop of the Oarai Gas Loop-1 and sweep gas of capsules. The temperature and burnup range of the in-pile experiments completed the HTTR fuel design limitations as shown in Fig. 3.4 [2]. These tests proved that there was no significant increase in additional fuel failure fraction up to the maximum burnup of the HTTR design.

3.1.2.5 Temperature limit of HTTR fuel The crystalline material comprising the SiC layer of the TRISO coating has a tendency to decompose at high temperature. The transition temperatures of β-SiC (asdeposited) to α-SiC vary from 1600 C to 2200 C [810]. The limitation of the fuel temperature of the HTTR was conservatively determined. In order to ensure fuel integrity at elevated temperatures, the behavior of the irradiated CFPs was examined in a temperature range up to 2400 C with a furnace installed in a hot cell [6]. A typical result of the accumulated failure fraction in the ramp tests is presented in Fig. 3.5 [2,9,11]. As shown in the figure, the failure begins at about 2200 C and almost 100% of the CFPs failed at 2400 C. From these tests, it was found that the coating layers of the HTTR CFPs would maintain their intactness below 1600 C. Thus the maximum

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185

1 GAT data (20°C/h,50°C/h,200°C/h) JAERI data (300°C/h)

Failure fraction

0.8

0.6

0.4

0.2

0 1000

1400

1800 2200 2600 Temperature (°C)

3000

Figure 3.5 Relation between failure fraction and fuel temperature [2].

fuel temperature shall not exceed 1600 C at any anticipated transient to avoid fuel failure, that is, to permit reutilization of the fuel after anticipated transients.

3.1.3 Fabrication technologies for HTTR fuel Fuel fabrication was undertaken with a laboratory scale capacity (10 kgU/y) at NFI, and a pilot plant with a capacity of about 40 kgU/y was subsequently developed in 1972. For irradiation experiments and out-of-pile characterizations, various fuels were produced by this pilot plant. The fabrication capacity was expanded to about 200 kgU/y in 1983 to produce the fuels for the Very High Temperature Reactor Critical Assembly (VHTRC). Also, the fuel elements for the Oarai Gas Loop-1 experiments carried out after 1984 were produced in this plant. The Oarai Gas Loop-1 fuels of last three generations, that is, fuels for the 13th to 15th Oarai Gas Loop-1 irradiation experiments, were fabricated using the coater with a mass production scale. By the commercial scale plant with the licensed capacity of 400 kgU/y launched at NFI in 1992, fabrications of the first and second loading fuels of the HTTR were carried out from June 1995 to December 1997 and from October 2002 to March 2005, respectively [1214].

3.1.3.1 Improvement of fuel fabrication process Fig. 3.6 depicts a flow diagram of the HTTR fuel production process, which consists of the fuel kernel, CFP, and fuel compact processes [1]. The UO2 kernels were fabricated in a gel-precipitation process. After formation of uranyl nitrate solution containing methanol and an additive, spherical droplets are produced by a vibration

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High Temperature Gas-cooled Reactors

Nitric acid

U3O8 powder

Additives

Uranyl nitrate (UO2(NO3)2)

UO2 kernel

Graphite powder

Binder

TRISO-CFP

Broth solution

1st layer (low-dense PyC) C2H2 (w/Ar) 2C + H2

Vibrational dropping into NH4OH

2nd layer (high-dense PyC) C3H6 (w/Ar) 3C + 3H2

Warm pressing

3rd layer (SiC) CH3SiCl3 (w/H2) SiC + 3HCl

Preliminary heat treatment

4th layer (high-dense PyC) C3H6 (w/Ar) 3C + 3H2

Heat treatment

ADU ((NH4 )2U2O7) particles Calcination UO3 particles

Graphite matrix

Overcoating

Reducing & sintering UO2 kernel

TRISO-CFP

Fuel compact

Figure 3.6 Flow diagram of the HTTR fuel production process [1].

dropping technique [13]. Following the drying and calcining, reduction of the calcined kernels to UO2 was carried out. Kernel fabrication was completed by a sintering process to produce dense UO2 kernels. Coating layers are deposited on the kernels in a chemical vapor deposition process using a fluidized bed type of coater. The TRISO-coating process is divided into four coating processes for the porous PyC, IPyC, SiC, and OPyC layers. The buffer and high-density PyC coating layers were derived from C2H2 and C3H6, respectively, and the SiC layer from CH3SiCl3. All UO2 kernels and TRISO-CFPs are classified by means of a vibrating table to exclude odd shape particles. The as-manufactured quality of the fuel has been improved by the modification of fabrication conditions and processes. The coating failure during coating process was mainly caused by the strong mechanical shocks to the particles given by violent particle fluidization in the coater and by the unloading procedure of the particles. The coating process was improved by optimizing the mode of the particle fluidization and by developing the process without unloading and loading of the particles at the intermediate coating process [15]. The fuel compacts of the HTTR are produced by warm pressing of the CFPs with graphite powder. First, natural graphite powder, electro-graphite powder, and a binder are mixed; then the mixture makes graphite matrix after fine grinding process. CFPs are overcoated with the graphite matrix and warm-pressed to make annular cylinder of green compacts. Green compacts are preliminarily heat treated for carbonization at 800 C under nitrogen atmosphere, then sintered at 1800 C under vacuum to make fuel compacts. The fabrication process was modified to reduce the defective particle fraction during the compaction process. The compaction process was improved by optimizing the combination of the pressing temperature and the pressing speed of the overcoated particles to avoid the direct contact with neighboring particles in the fuel compact [1]. In the beginning of the first-loading fuel fabrication of the HTTR, unexpected large SiC-failure fractions, about 35 particles in a fuel compact, were observed. Then, relations between the measured SiC-failure fractions and fabrication

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parameters, such as coating layer thickness of the CFP, overcoating layer thickness, and pressing speed, were analyzed during fabrication. Finally, it was concluded that the SiC layer thickness should be thicker than 27 μm to avoid as-fabricated SiC-failure during the compaction process. In addition, odd-shaped overcoated particles were removed before compaction process. After these improvements, a significant SiC-failure was no longer observed during fabrication [12]. In each campaign of the HTTR fuel fabrication with optimized fabrication process mentioned above, 0.9 ton of uranium was used, and 66,780 fuel compacts corresponding to 4770 fuel rods for 150 fuel assemblies were successfully produced. As a result, as-fabricated fuel compacts contained almost no through-coatings failed particles and few SiC-defective particles as shown in Fig. 3.7. The figure shows histograms of the presence of the fuel compact containing through-coatings failed or SiC-defective particles. Abscissa of Fig. 3.7 shows the number of through-coatings failed or SiC-defective particles per one fuel compact, which were measured by acid-leaching or burn-leaching methods, respectively [12]. Fuel compacts have been fabricated with 126 and 128 fuel compact batches of the first and second loading fuels, respectively. Through-coatings and SiC-defective fractions were measured with two and three fuel compacts sampled from one fuel compact batch (i.e., 252/256 and 378/384 fuel compacts were used for through-coatings failure/SiC-defective fractions of the first-and second loading fuel, respectively). Average through-coatings and SiC defective fractions for the first and second loading fuels were 2 3 1026 and 8 3 1025, and 8 3 1025, and 1.7 3 1024, respectively [12,14]. It was concluded that these values were quite lower than the criteria, 1.5 3 1024 for through-coatings failure and 1.5 3 1023 for SiC-defective fraction, and the HTTR fuel has been fabricated successfully with high quality. 100% 80% 60%

1st loading fuel Through-coatings failure SiC-failure

Frequency

40% 20% 0% 100%

80%

2nd loading fuel Through-coatings failure

60%

SiC-failure

40% 20% 0%

0 1 2 3 4 5 6 7 8 9 10 Number of failed particles in a fuel compact

Figure 3.7 Number of failed particles in fuel compact in first and second loading fuel fabrications [14].

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High Temperature Gas-cooled Reactors

3.1.3.2 Quality control The inspection items for the HTTR fuel were determined to confirm specifications, which certify nuclear and thermalhydraulic design, irradiation performance, and so on. From the viewpoint of purposes, the inspection items are divided into three categories, namely (1) compulsory, (2) user’s requirement or optional, and (3) vender’s quality control [4]. The sampling rate was also determined by considering the uniformity of inspected data. Three categories are basically classified as (1) smallscattering data, (2) medium-scattering data, and (3) large-scattering data. One sample is measured from an inspection lot for the small-scattering data. For the inspection lot with medium-scattering data, three samples are measured and all of them should satisfy criterion. For the large-scattering data, measured data should meet statistically required criterion with 95% confidence. The inspection item, purpose, method, and sampling rate in the HTTR fuel fabrication are summarized in Table 3.2 [4]. JAEA and Japanese manufacturers have developed HTGR technologies to achieve those criteria, and own the following patents, know-hows, and data [16]. Table 3.2 Inspection item, purpose, method, and sampling rate in HTTR fuel fabrication [4]. Inspection item

Purposea

Method

Sampling rate

B

Mass spectrometer analysis and gamma-ray spectrometer analysis Optical particle size analysis Optical particle size analysis Mercury substitution Oxidation and weighing Emission spectrometer analysis

1 sample/enrichment

Fuel kernel 235

U enrichment

Diameter

B

Sphericity

A

Density O/U ratio Impurities

B A A, B

1 sample (100 kernels)/ fuel kernel lot 3 samples (100 kernels/ sample)/fuel kernel lot 3 samples/fuel kernel lot 1 sample/fuel kernel lot 1 sample/enrichment

CFP Layer thickness

A

Solvent substitution or sink-float Polarization photometer

Optical anisotropy factor Diameter

A

B

Appearance

A

Optical particle size analysis Visual observation

Cross-section

A

Ceramography

3 samples/CFP lot 1 sample (5 CFPs/ sample)/enrichment 1 sample (100 CFPs)/CFP lot 1 sample (2000 CFPs)/ CFP lot 1 sample (20 CFPs)/CFP lot (Continued)

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Table 3.2 (Continued) Inspection item

Purposea

Method

Sampling rate

Sphericity Strength

A A

Selection by vibration table Point crushing

All CFPs 30 CFPs/enrichment

D

Mass spectrometer analysis and gamma-ray spectrometer analysis Gamma-ray spectrometer analysis Oxidation and weighing Density, impurities, grain size, and water content Contents, ash, melting point, and impurities Deconsolidation and acid leaching Burn and acid leaching

1 sample/enrichment

Fuel compact 235

U enrichment

U content

B

O/U ratio Graphite powder Binder

A A

Exposed U fraction SiC-failure fraction Packing fraction Matrix density Dimensions Appearance Marking Strength Cross-section Impurities a

A A A B A C A D A A B

Weighing and calculation Weighing and calculation Micrometer Visual observation Visual observation Compression Ceramography Emission spectrometer analysis

All fuel compacts 1 sample/compact lot 1 sample/graphite powder lot 1 sample/binder lot 2 samples/compact lot 3 samples/compact lot 3 samples/compact lot 3 samples/compact lot All compacts All compacts All compacts 3 samples/enrichment 1 sample/compact lot 1 sample/enrichment

A: Irradiation performance, B: Nuclear design, C: Thermalhydraulic design, D: Process control.

3.1.4 Performance of HTTR fuel during long-term high temperature operation In order to confirm the fuel integrity under severe irradiation conditions during the long-term high temperature operation of the HTTR, primary coolant sampling measurements were carried out. As a result, concentrations of fission gas nuclides of 85mKr, 87 Kr, 88Kr, 133Xe, 135Xe, 135mXe, and 138Xe were less than 0.1 MBq/m3. Fig. 3.8 shows release-to-birth rate ratio (R/B) of 88Kr during this operation [14]. Ranges of (R/B)s of German and US are referred from the references [17,18]. Measured (R/B) values resulted less than 1.2 3 1028, how was by four orders of magnitude lower than the design limit for the normal operation, 1 3 1024, corresponding to 0.2% of fuel failure. Fig. 3.9 shows (R/B)s of 88Kr as a function of the reactor power [1]. The (R/B) values are as low as 2 3 1029 up to 60% of the reactor power, then increase to 7 3 1029 at full power operation. During the high temperature test operation, where

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High Temperature Gas-cooled Reactors

10–4

10–5

(R/B) of 88Kr

(USS) FSV 10–6 (Germany) AVR 10–7

10–8 Long term high temperature operation 10–9

4- 11 18 25 1- 8- 15 22 1- 81 2 Ja -J -J -J Fe Fe -F -F Fe M 5-M 2-M n ar an an an b b eb eb b ar ar

Operation date

Figure 3.8 Release-to-birth rate ratio (R/B) of 88Kr during long-term high temperature operation of HTTR [14]. Reactor power (%) (operation mode:950°C/850°C) 0

20

40

60

80

100

15 20

40

60

80

100

(R/B) of 88Kr (10–9)

Continuous high temperature operation 10

5

Previous 950°C operation 850°C operation Predicted (sum of (a) to (d))

(a) (b) (c)

0 200

400 600 800 Reactor outlet coolant temperature (°C)

(d) 1000

Figure 3.9 (R/B)s in HTTR during 950 C and 850 C operations; (a) Diffusional release from contaminated uranium; (b) Recoil release from contaminated uranium; (c) Diffusional release from through-coatings failed particle; (d) Recoil release from through-coatings failed particle [1].

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the outlet coolant temperature is 950 C, the (R/B) became 1.5 3 1028 at full reactor power. The obtained data were analyzed by fission gas release model [19], and the fission gas release mechanism is recoiled from the contaminated uranium in the fuel compact matrix in lower reactor power. Beyond 60% of the reactor power, fractional release increases because diffusion release becomes main release mechanism. The increase of (R/B) in the high temperature operation is caused by increase of diffusion release according to fuel temperature elevation [19]. Finally, it was concluded that the HTTR fuel performed excellent quality during the long-term high temperature operation. The burnup in the end of this operation has reached around 11 GWd/t, although a little increase of the fission gas release was observed. It was suggested that the gap generating between fuel compact and graphite sleeve according to irradiation-induced shrinkage of fuel compact [6] would occur to make fuel temperature increasing [1].

3.2

Core components and reactor internals

3.2.1 Introduction The 30-MWt, seven-region core of the HTTR consists of hundreds of pin-in-block type fuel elements and is supported by a CBS. Helium gas enters the core after reversing at the upper plenum in the reactor vessel. Then the helium gas is heated up to 950 C flowing downward through fuel elements in each core region and leaves the reactor, passing through an inner tube of the outlet gas duct after mixing in the hot plenum. The HENDEL was constructed to investigate the thermal and hydraulic performance and structural integrity of the core and CBS. The HENDEL was composed of helium gas circulation loops, Mother (M1 and M2), and Adaptor (A), and two test sections, that is, the fuel stack test section (T1) and in-core structure test section (T2) as shown in Fig. 3.10. The T1 test section consisted of a single-channel test rig (T1-S), which is a mock-up model of a single fuel element and a multichannel test rig (T1-M), which is a mock-up model of fuel elements, namely a fuel column. The basic thermal and hydraulic characteristics of the fuel element were investigated with T1-S simulating heat flux conditions of fuel rod such as uniform, exponential, and cosine distributions in axial direction [20,21]. The influence of power unbalance among fuel elements on the thermal and hydraulic characteristics was investigated with T1-M [20,22]. The T2 test section was a full-scale mock-up model of the CBS. The thermal and hydraulic tests were performed to verify the following items under the HTTR conditions [20]: (1) the sealing performance of helium gas among permanent reflector blocks [23], (2) the mixing performance of helium gas in the CBS [24], (3) the insulation performance of the CBS [20], and (4) the thermal performance of a coaxial hot gas duct [25].

3.2.2 Tests on core components 3.2.2.1 Test apparatus The T1-S rig provided a simulated fuel channel for testing simulated fuel rod. Helium gas flowed downward through the annular channel between the fuel rod

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High Temperature Gas-cooled Reactors

Figure 3.10 Schematic flow diagram of HENDEL loop [20].

and an outer tube and was heated up to 1000 C. The simulated fuel rod consisted of seven subrods and upper and lower electrodes. The dimensions of each subrod were 570 mm in total length, 460 mm in effective heating length, and 46 mm in outer diameter. Nine spacer ribs were machined on the outer surface of each subrod to maintain an annular gap between the simulated fuel rod and the outer tube. The dimensions of the outer tube of Alloy 800 were 53 mm in inner diameter and 6370 mm in length. Nine thermocouples and ten pressure taps were embedded in the tube wall to measure wall temperatures and the pressure drop of helium gas in the channel. The T1-M rig simulated one column of the stack of fuel block. The simulated fuel block was a hexagonal block of 570 mm in height and 299 mm in width across the flats. Eleven simulated fuel blocks were arranged in one line and twelve simulated fuel rods were installed in each block as shown in Fig. 3.11. Helium gas temperatures were measured at the inlet and outlet of each flow channel and local temperatures were obtained by interpolation. The heating surface temperatures were measured at 410 mm from top of each subrod. Temperature distributions of the simulated fuel blocks were also measured.

3.2.2.2 Thermal hydraulic characteristics The thermal hydraulic characteristics of fuel rod on pressure drop and heat transfer were investigated with the T1-S rig.

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Figure 3.11 Schematic view of multichannel test rig (T1-M) [22].

3.2.2.2.1 Pressure drop The pressure drop in the fuel channel is defined as follows: Δp 5

2ρu2 ; De f

(3.1)

where Δp, ρ, u, De, and f are the pressure drop, the density, the velocity, the hydraulic diameter, and the friction factor, respectively. The obtained friction factors were compared with those for the parallel plates and the Blasius equation. The obtained friction factors were correlated by the following equations: f5

28 ðRe . 2000Þ; Re

f 5 0:094Re20:25 ðRe $ 2000Þ:

(3.2) (3.3)

The friction factors were 17%19% higher than those of the parallel plates and the Blasius equation. It appears that the friction factors increased due to the effects

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High Temperature Gas-cooled Reactors

of turbulence promoted by the spacer ribs and the velocity change resulting from the changes in the cross-sectional area of the flow channel.

3.2.2.2.2 Heat transfer Fig. 3.12 shows the relationship between the Nusselt and the Reynolds numbers obtained from the tests. These results were calculated by hw 5

qc ; Tw 2 Tg

(3.4)

Nu 5

hw De : λg

(3.5)

The numerator in Eq. (3.4) is the convective heat flux calculated by  4  σ Twc 2 Tb4  ; qc 5 qe 2 qr 5 qe 2 1=εw 1 Di =Do 3 ε1b 2 1

(3.6)

where h, q, T, λ, ε, and σ are the heat transfer coefficient, the heat flux, the temperature, the thermal conductivity, the emissivity, and StefanBoltzmann constant, and the subscripts b, c, e, g, i, o, r, and w are the outer wall of fuel channel, the convective heat transfer, Joule heating, the bulk temperature of helium gas, the inner and outer of fuel channel, the radiative heat transfer, and inner wall of fuel channel, respectively.

Figure 3.12 Relationship between Nusselt number and Reynolds number [20].

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The last term on the right-hand side of Eq. (3.6) represents the radiative heat flux, where εw(50.8) and εb (50.3) were obtained by the other component test. The Twc is the surface temperature at the center of each subrod, determined by using least square interpolation. The solid lines in Fig. 3.12 show the Nusselt numbers for a concentric smooth annular channel defined by Kays and Crawford [26] and Dalle Donne and Meerwald [27]. The dashed line shows the Nusselt numbers obtained from the tests under conditions of uniform, exponential, and cosine heat flux distributions. The line is defined by Nu 5 5:6ðRe # 2700Þ;

(3.7)

Nu 5 0:0215Re0:8 Pr0:4 ðRe $ 2000Þ:

(3.8)

The results agreed with Eqs. (3.7) and (3.8) within 6 10%, and no rapid decrease in the Nusselt number over the traditional region was observed. In the turbulent and laminar regions, the Nusselt numbers were about 17%21% higher than those of a smooth annular channel due to the effects of turbulence by the spacer ribs and the velocity change resulting from the changes in the cross-sectional area of the flow channel.

3.2.2.3 Effect of power unbalance The effect of power unbalance among the simulated fuel rods on flow and temperature distributions was investigated with the T1-M rig. In the test, the uneven power distribution unbalance was given from the 11th to 8th among the simulated fuel rods as shown in Fig. 3.13, because the power distribution caused the maximum temperature distribution in the plane of the simulated fuel block. The maximum deviation of the power from the mean value was 6 10%. In the figure, the solid and dasheddotted lines mean the calculation results based on measured temperatures. The

Figure 3.13 Axial temperature distribution obtained by uneven power distribution test [20].

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High Temperature Gas-cooled Reactors

maximum deviations of the heating surface temperature and outlet helium gas temperature were about 100 C and 20 C, respectively. It was obvious that the thermal conduction through simulated fuel block flattened the temperature distribution in the radial direction. With increasing of the Reynolds number, the deviation of flow rate in the channel decreased. The maximum deviation of the flow rates was about 4% of the mean value, which was observed for low Reynolds number condition of 3000.

3.2.3 Tests on reactor internals 3.2.3.1 Test apparatus The T2 test section was a full-scale mock-up model of the CBS of HTTR. The CBS to be tested, flow regulations, region heaters, an outlet hot gas duct, and the other auxiliary equipment were contained within a pressure vessel, as shown in Fig. 3.14.

Figure 3.14 Schematic view of in-core structure test section (T2) [20].

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Figure 3.15 Schematic view of HTTR core bottom structure [24].

There were inlet and outlet passages for both hot and cold helium gas between the M2 1 A section and the T2 test section. In the T2 test section, the cold helium gas of 400 C was supplied from the M2 1 A section, flowed into the pressure vessel through the outer tube of the outlet hot gas duct, was guided upward along the outer surface of the permanent reflector blocks, and returned to the M2 1 A section through the outer tube of the inlet gas concentric duct. Hot helium gas, heated to 930 C by the M2 1 A section, entered the T2 test section through the inner tube of the inlet concentric duct and flowed through the regulators and region heaters to the CBS. Fig. 3.15 shows the general view of the CBS of the HTTR composed of hot plenum blocks, core support posts, lower plenum blocks, carbon blocks for insulation, a core support plate, an outlet gas duct, etc. The cold helium gas at 400 C enters in the pressure vessel by way of an outer tube of the outlet gas duct and flows upward along the outer surface of the permanent reflector blocks. The hot helium gas at 950 C heated in the core flows downward in the CBS and leaves by way of an inner tube of the outlet gas duct.

3.2.3.2 Sealing performance of helium gas The sealing performance between the permanent reflector blocks is essential for the HTTR to achieve the core outlet gas temperature of 950 C. If the leakage flow, due to inaccuracy of fabrication or other reasons, exceeds a limit value, the reactor coolant flow must be increased to avoid damage to the fuel. As a result, if the leakage were severe, it would be impossible to achieve the core outlet gas temperature of 950 C. The leakage flow rate depends mainly on the size of the gaps between adjacent permanent reflector blocks and the pressure difference across them. Sealing performance tests were performed with various pressure differences. The results of a total function test performed immediately after the construction of the T2 test section revealed that the average gaps were about one fifth the anticipated design value of 0.2 mm. Most of the exiting leakage flow was determined to be

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High Temperature Gas-cooled Reactors

turbulent flow through the corner conduits between the permanent reflector blocks rather than laminar flow through the gaps between them. This is because the gap caused by the corner conduits was larger than other gaps. The corner conduits arose from a 2 to 3 mm scraping of the corner edges to prevent cracking from the sharpened edge during transportation and fabrication. The results also showed that the leakage flow rate was affected by the condition of the cold helium gas. Leakage flow rates under different conditions may be calculated using reference data and the following equation:  4=7  1=7    W ρ μR ρ μR 5X 1 ð1 2 X Þ WR ρR ρR μ μ

(3.9)

where W, X, and μ are the leakage flow rate, the fraction of turbulent flow to the total leakage flow, and the viscosity, the subscript R is the reference condition, respectively. The first term on the right-hand side of Eq. (3.9) represents the turbulent leakage flow along the corner conduits between the permanent reflector blocks, and the second term represents the laminar leakage flow through the gaps between them. Fig. 3.16 shows the relationship of the leakage flow rate to the pressure difference inside and outside the permanent reflector block at a temperature of 400 C and a pressure of 4.0 and 2.0 MPa. The solid line indicates the most probable values

Figure 3.16 Relationship between leakage flow and pressure difference between inside and outside of permanent reflector blocks [20].

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from the total function test (400 C, 4.0 MPa), and the dashed line indicates the calculated values at 400 C and 2.0 MPa using Eq. (3.9) and the test data at 4.0 MPa and 400 C. The calculated values agreed with the measured data at the same conditions to within 8%. The figure also compared the initial data (total function test) and recent data (data after 4000 h of operation) at the same conditions of temperature (400 C) and pressure (4.0 MPa). No detectable change in the leakage flow rate has occurred since T2 test section was constructed. This indicates that there is no enlargement of the gaps induced by irregular dimensional changes of the blocks or by degradation of the core restraint band mechanism for the permanent reflector blocks.

3.2.3.3 Mixing performance of helium gas The HTTR core is divided into seven regions, each composed of stacked fuel elements. Heated helium gas from each region is mixed in a hot plenum located at the CBS. If the helium gas is not mixed to a uniform temperature in the hot plenum, the temperature differences in helium gas flow may cause hot spots in the high temperature components, such as an IHX and a primary water cooler (PWC), and inaccurate measurement of the core outlet gas temperature required to determine the thermal output of the HTTR. Thermal mixing tests were conducted using the T2 test section to examine the mixing characteristics of the hot helium in the HTTR. The helium gas temperatures were measured at the hot plenum blocks, hot plenum, and outlet gas duct at the location L/Di 5 5.5 and 12.5. The low-temperature test was performed with both the cold and hot helium gases at 300 C to clarify only the thermal mixing characteristics in the CBS. The high temperature test was carried out with the cold helium gas at 400 C and the hot one at 900 C so that both the mixing and the heat loss of the hot helium were considered. The temperature of the hot helium gas in some regions was higher than that in others using the region heaters. The temperature difference between hot and cold regions ranged from 40 C to 100 C, and the Reynolds numbers in the outlet gas duct were from 1.8 to 4.7 3 105. The helium gas was mixed sufficiently in the hot plenum, and there was no hot streak in the outlet gas duct when the helium gas temperature in the central region was high. However, the mixing of the hot helium gas was insufficient in the hot plenum when the helium gas temperatures in some circumferential regions are higher, and a hot streak occurred along the wall of the outlet gas duct as shown in Fig. 3.17. In the figure, symbols means obtained temperature at two hot circumferential regions in the tests, and solid lines are analytical results. A dimensionless temperature difference, Δθmax, is defined by the following equation using the temperature difference between hot and cold regions, ΔTp, and the helium gas temperature in cold region, Tp,min: Δθmax 5

T 2 Tp;min : ΔTp

(3.10)

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High Temperature Gas-cooled Reactors

Figure 3.17 Axial variation of radial temperature distribution in outlet gas duct in case of hot circumferential region [24].

Here, one hot region means that the helium gas temperature is higher in only one circumferential region, and two and three hot regions indicate that the helium gas temperature in the adjacent circumferential regions was higher, respectively. The temperature difference at L/Di 5 20 was approximately 10% of ΔTp in the case of three hot regions. Considering the locations of the IHX and the PWC of the HTTR, the measuring position of the reactor outlet temperature of the helium gas installed at L/Di . 30 from the core, and the temperature difference among the circumferential regions (ΔTp 5 20 C), we can conclude that the hot streak has little effect on the IHX and the PWC.

3.2.3.4 Insulation performance of core bottom structure An insulation performance test was conducted for hot helium gas having a temperature of 900 C1000 C, a flow rate of 3.5 kg/s, and a pressure of 4.0 MPa. The temperature, flow rate, and pressure of the cold helium gas were 400 C, 4.0 kg/s, and 4.0 MPa, respectively. Fig. 3.18 shows temperature distribution of the CBS with a uniform temperature distribution of 950 C hot helium gas. The abscissa displays the region number and orientation angle. The temperature distribution at each longitudinal cross-section was uniform in the circumferential direction. The temperatures of the metal structures such as the core support plate and were lower than the expected design temperatures. This occurred because the metal structures were adequately insulated by the thermal insulation blocks and cooled by the cold helium gas.

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Figure 3.18 Temperature distribution of components in core bottom structure [20]. (A) Direction parallel to hot He gas flow. (B) Direction perpendicular to hot He gas flow.

3.2.3.5 Thermal performance of a coaxial hot gas duct A coaxial hot gas duct with an internal insulator, where hot helium gas of 950 C flows in the inner tube and cold helium gas of 400 C in the outer one, is used for the primary and the secondary coolant system of the HTTR. There are a number of studies on the thermal performance of the single hot gas duct presented by Bro¨ckerhoff [28], Hishida et al. [29], etc. However, there are few studies on the coaxial hot gas duct. An experimental study on the thermal insulation performance of the coaxial hot gas duct was carried out with the T2 test section [25]. The structural integrity was also investigated on a part where the coaxial hot gas duct is separated into two single hot gas ducts for hot and cold helium gases. Fig. 3.19 shows the structure of the coaxial hot gas duct vertically installed in the CBS of the T2 test section. The internal insulation layer is installed between the liner and inner tubes to assure temperature of the inner tube below design value and to prevent heat exchange between hot and cold helium gases. The thermal insulation performance was investigated on the coaxial hot gas duct by measuring temperature distributions of hot and cold helium gases, and the liner and inner tube surfaces with thermocouples in the following conditions: a hot helium temperature of 450 C950 C, a flow rate of 2.04.0 kg/s, and a pressure of 1.04.0 MPa, and a cold helium temperature of 50 C400 C, a flow rate of 1.04.0 kg/s, and pressure of 1.04.0 MPa. The structural integrity of the part where the coaxial hot gas duct is separated into two single hot gas ducts was also investigated on the coaxial hot gas duct B by measuring the temperature distribution of the inner and outer tube surfaces with thermocouples in the following conditions: a hot helium temperature of 530 C730 C, a flow rate of 1.53.0 kg/s, and pressure of 3.0 MPa, and a cold helium temperature of 170 C200 C, a flow rate of 1.53.0 kg/s, and a pressure of 3.0 MPa.

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High Temperature Gas-cooled Reactors

Figure 3.19 Cross-section of coaxial hot gas duct [25].

A relationship between the mean temperature and the effective thermal conductivity of the thermal insulator in the coaxial hot gas duct A is shown in Fig. 3.20. The effective thermal conductivity of the single hot gas duct is also shown [29]. Both the effective thermal conductivity of the coaxial and the single hot gas ducts are nearly equal. Moreover, there is no difference on the effective thermal

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Figure 3.20 Relation between effective thermal conductivity and mean temperature of thermal insulator of coaxial hot gas duct [25].

conductivity within the helium gas pressure range of 1.04.0 MPa. This result verified that the effect of natural convection in the thermal insulator could be safely ignored. The obtained effective thermal conductivities of the thermal insulator, λeff, were fitted by a linear form as follows: λeff 5 2:49 3 1024 1 0:34ðW=mKÞ:

3.3

(3.11)

Passive cooling system

3.3.1 Introduction Modular high temperature gas-cooled reactors (MHTGRs) are advanced nuclear reactors under development that employ passive cooling to enhance its inherent safety. The passive cooling system of the MHTGR removes decay heat by means of thermal conduction, radiation, and natural convection sufficiently so as to protect the reactor pressure vessel from overheating and to limit nuclear fuel temperature so as to prevent release of fission products during accidents. A few kinds of passive cooling systems have been proposed, including a forced water-cooling panel system of HTTR [2], in which decay heat is removed indirectly from the pressure vessel by forced circulating water in cooling pipes surrounding the pressure vessel in a reactor cavity filled with air. However, hot spots could appear on important components such as the pressure vessel superheated by a natural convection flow of the superheated gas, which could cause damage of the components in the system during accident conditions. On the

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other hand, some HTGR designs among others have multiple standpipes installed on the upper head of the pressure vessel. These standpipes appear in actual reactors to house the control rods and the control rod driving mechanisms. The standpipes reduce the heat transfer area of the upper head and restrict thermal radiation from the upper head to the cooling panel. This could cause the onset of hot spots on the upper head of the pressure vessel during accident conditions. However, the threedimensional structure of the upper head of the pressure vessel with many standpipes makes the evaluation of its temperature distribution difficult. Thus in order to design water-cooling panel system, it is important to establish suitable design and evaluation methods for the system so as to properly evaluate the effects of natural convective flow of superheated gas on the temperature distribution of such critical components as the pressure vessel, especially in the vicinity of the standpipes. The experiments were carried out using an experimental apparatus to evaluate the temperature distributions of the components and the heat removal performance of the water-cooling panel system [30]. The experimental apparatus consisted of a pressure vessel of 1 m in diameter and 3 m in height, which contained internal heaters with a maximum heating power of 100 kW to simulate the decay heat of the core, and was surrounded by external cooling panels, which simulated the water-cooling panel system of HTTR. The experiments performed in the pressure vessels, without and with standpipes on the upper closure head, were compared to assess the effects of the standpipes on both heat removal performance and temperature distribution of the water-cooling panel system. The numerical code, THANPACST2 [31], which solved the flow and temperature distributions of the components simultaneously by considering heat transfer through thermal conduction, radiation, and natural convection, was developed to assist in the design and evaluation of an HTGR passive cooling system. In contrast to other earlier codes [32], the THANPACST2 code is able to consider the effects of natural convection of the superheated gas on the heat transfer performance and temperature distribution of the components. A new axisymmetric model has been built into the axisymmetric THANPACST2 code to simulate the three-dimensional configuration of the upper head of the pressure vessel by using porous body cells, which have sets of the characteristics of fluid and solid. The experimental data were used to validate the numerical method, models, and thermal properties of the components. Hence, the simulation was carried out to evaluate the effects of such natural convection with respect to the temperature distributions of the components and the heat removal performance of the system. The simulation showed that the thermal radiation was responsible for the majority from 70% to 86%, of the total heat transferred to the simulated cooling panel, and that the ratios of the heat transferred with natural convection in the reactor cavity to the total heat transferred to the cooling panels were varied. The experimental results were provided for code-experiment benchmark problems to the IAEA Coordinated Research Project on “Heat Transport and Afterheat Removal for Gas Cooled Reactors Under Accident Conditions” [33]. By comparing the numerical results of the THANPACST2 code with the safety evaluation codes, the effects of such natural convection with respect to the temperature distributions of the components and the heat removal performance of the system were revealed.

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3.3.2 Experiment 3.3.2.1 Experimental apparatus A schematic of the experimental apparatus simulating the cooling panel outside the pressure vessel is shown in Fig. 3.21. The experimental apparatus consists of a pressure vessel, which contains internal heaters with a maximum heating power of 100 kW to simulate the decay heat of the core, and is surrounded by external cooling panels in the reactor cavity occupied by air. The height, diameter, and thickness of the pressure vessel made of stainless steel are 3, 1, and 0.012 m, respectively. The upper head is connected with the shell of the pressure vessel by a flange. The pressure vessel is supported by a simulated skirt-type support, which has airflow channels including the rectangular gap between the H-type leg and the pressure vessel, and the gap between the lower edge of the skirt and the ground. Nineteen simulated standpipes are set up on the top of the upper head of the pressure vessel systematically as shown in Fig. 3.22. The simulated standpipes are detachable. The simulated standpipe consists of stainless steel pipe filled with thermal insulation of KAOWOOL (ceramic fiber insulation) since thermal insulation and radiation shield are located between the control rod mechanism and reactor core inside the standpipe of an actual HTGR. The upper ends of the standpipes are uniformly located 330 mm above the upper head of the pressure vessel. The cooling panels are installed individually above, below, and around the side of the pressure vessel. The numbers of the cooling pipes are 22, 12, and 88 for the upper, lower, and side cooling panels, respectively. Both ends of the cooling pipes are connected to ring type headers. Black heat-resistant paint is coated on the outer surfaces of the pressure vessel, the cooling panels, and the standpipes to obtain an emissivity of almost one. Water is supplied with the maximum flow rate around 10 m3/h. Thermal insulation made of KAOWOOL is applied on the outside of the cooling pipes to avoid any external disturbance. The inner diameter, height, and thickness of the thermal insulation are 2, 4, and 0.1 m, respectively. The helium gas pressure is varied within an upper limit of 1.1 MPa in the pressure vessel. The diameter and height of the simulated core are 0.6 and 2 m, respectively. The maximum temperature of the simulated core is 600 C. The simulated core is divided into six heater segments. Lateral surfaces of Nos. 25 segments are heated with Nichrome helical coils wound around annular-type ceramic blocks filled with thermal insulation. Numbers 1 and 6 segments on the top and bottom, respectively, are heaters with Nichrome coils fixed on the horizontal surfaces of dish-type ceramic blocks. The heater segments are held by two annular plates mutually braced by four stainless steel rods at the top and bottom. The bottom plate is in turn supported by eight legs on the lower head of the pressure vessel.

3.3.2.2 Measurement Temperatures of the surfaces of the pressure vessel, simulated core, cooling panel tubes, and thermal insulation are measured by sheathed chromelalumel thermocouples. Temperatures of the outer surface of the pressure vessel are measured by

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Figure 3.21 Schematic view of experimental apparatus [30].

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Figure 3.22 Configurations of standpipes on upper head of pressure vessel [30].

155 thermocouples at four different angles with 90-degree intervals. The number of thermocouples attached on the surfaces of the simulated core, the cooling pipes, and the thermal insulation are 22, 12, and 17, respectively. Temperatures of the helium gas are measured by three thermocouples with radiation shielding plates at a distance of 50 mm from the inner surface of the upper head of the pressure vessel. The cooling water temperature is measured by a platinum resistance bulb at the inlet and outlet of each water line of the cooling panels. The flow rate of the cooling water is measured by a magnetic flow meter. The heating value for each heater segment of the simulated core is measured by a power transducer, and the total measured heating rate is compared with the total measured heating transferred to the cooling panels.

3.3.2.3 Experimental conditions Varying fluid temperature and pressure inside the pressure vessel as well as experimental set-ups with and without standpipes on the top of the pressure vessel, three cases of experiments were compared to investigate the heat removal performance of the water-cooling panel system. The temperatures of the components, the heat input of the simulated core, and the heat transferred to the cooling panels are measured. Experimental conditions are summarized in Table 3.3. In Case 1, helium gas is evacuated from the pressure vessel. In Case 2, the pressure vessel was pressurized by helium gas at 0.73 MPa. In these cases, the temperatures of thermocouples for power control of the heater segments of the simulated core are 300 C. The effects of natural convection of helium gas on the temperature distribution of the components and the heat removal performance of the system are evaluated

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Table 3.3 Experimental conditions [30]. Case number 1

2

3

1.3 3 1026

0.73

0.43

13.14

28.79

77.54

1.01 2.31 2.64 2.46 3.76 0.96 Water Without

1.16 3.11 3.52 5.10 10.42 5.49 Water Without

5.63 19.60 21.59 22.70 0.00 8.00 Water With

Gas Pressure inside pressure vessel (MPa) Heater input (kW) Total input Heater segment No. 1 No. 2 No. 3 No. 4 No. 5 No. 6 Cool line panel type Standpipe

by comparing the experimental data and the numerical results. In Case 3, the effect of natural convection of superheated gas and the effect of the standpipes on restricting heat transfer to the water-cooling panels are evaluated. The temperatures of thermocouples for power control of the heater segments of the simulated core are 600 C. In this case, helium gas pressure inside the pressure vessel is lowered and the heater segment No. 5 of the simulated core is powered off so that the maximum allowable temperature and stress of the pressure vessel are kept lower than the respective maximum allowable limits of the pressure vessel steel.

3.3.3 Numerical method 3.3.3.1 Numerical code Governing equations in the THANPACST2 code are NavierStokes equations of continuity, momentum, and energy for incompressible fluid and heat transfer for solid as follows: Continuity: r~ u 5 0:

(3.12)

Momentum:     @~ u 1 1 ρ 2 ρ0 2 5 2 rP 2 r ζ j~ u j 1 rðνr  ~ u Þ 2 ð~ u  rÞ~ u 2~ g : @t ρ0 2 ρ0

(3.13)

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Energy: @T 1 ½r  ðkrT Þ 1 qI 1 qS  2 ð~ 5 u  rÞT: @t ρ0 C P

(3.14)

Heat transfer: @TS 1 ½r  ðkS rTS Þ 2 qI 2 qS 1 qrad 1 qP : 5 ρS C @t

(3.15)

where kS, qI, and qS are thermal conductivity of solid, heat transferred from solid to fluid in a porous body cell, and heat transferred from solid to fluid, respectively. Eqs. (3.12) and (3.13) are employed for simplicity of calculation in this method although the equations do not reflect the exact characteristics of the gas experiencing variable thermal properties in a finite volume. The buoyancy force generated by the spatial difference of fluid temperature, the fifth term on the right side of Eq. (3.13), is calculated using Boussinesq’s approximation. Governing equations of fluid and solid (3.12)(3.15) are discretized together in a porous body cell [34]. Geometric factors to calculate qrad in Eq. (3.15) are obtained by the unit-sphere method [35]. The sum of the geometric factors on the surface of a cell is one. The convergence of calculation is judged by the criterion that the heat balance is within an accuracy of 1% between the heat transmitted from the surfaces, such as the simulated core and the pressure vessel, and that transferred to the surfaces of the cooling panels.

3.3.3.2 Numerical model Fig. 3.23 shows the grid system of the experimental apparatus. As seen, an axisymmetric cylinder of a 23 3 39 stagger grid geometry is used to discretize the simulated core, the pressure vessel with the standpipes on the top of the pressure vessel, the cooling panels, and the simulated skirt-type support with air channels at the upper and lower edges in the numerical model. In addition, the axisymmetrical code assumes uniform surfaces for the simulated core, the cooling panels, and H-type legs of the simulated skirt-type support in spite of their three-dimensional structures. The air flow channels of the simulated skirt-type support are assumed to be axisymmetry. The three-dimensional configuration of the 19 standpipes is modeled by two layers of concentric cylinders surrounding a center cylinder for the simulations. To consider the effects of the airflow through the standpipes, the two outer layers of concentric cylinders are modeled by porous body cells, as shown in Fig. 3.24. Thermal conductivity of the standpipes is assumed to be the value of thermal insulation material, which is considerably lower than that of the standpipe. Finally, slip condition was assumed in the velocity boundaries on the walls of components. Thermal radiation heat transfer is calculated between the inner surface of the pressure vessel and the simulated core, between the pressure vessel and the

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Figure 3.23 Differential scheme for numerical analysis [30].

simulated skirt-type support and the cooling panels, and among the standpipes. The cooling panels are composed of two heat sinks; they are the cooling pipes and the outermost thermal insulation of the cavity. Although some heat is lost through the outermost thermal insulation, it is minimal and thus neglected in the study. Furthermore, the reflective thermal radiation from the inner surface of the thermal insulation to the pipes of the cooling panels is also insignificant and thus neglected. Therefore the total heating value of the simulated core is assumed to be transferred only to the cooling panels. Emissivity of the black paint coated on the surfaces of the pressure vessel, skirttype support, standpipes, and cooling panels is assumed to be 0.95, which was found to be a suitable value in a preliminary evaluation [31]. Emissivity is assumed to be 0.79 for the inner surface of the pressure vessel, on which no coating is applied, and 0.7 for the Nichrome heating surface of the simulated core, both of which are given in the literature [36]. Thermal radiation is calculated between one solid and a maximum of 30 neighboring solid cells.

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Figure 3.24 Porous body cell model simulating standpipes [30].

3.3.3.3 Empirical correlations for natural convective heat transfer The heat transfer rate used to calculate qI and qS in Eqs. (3.14) and (3.15) is computed based on measured temperatures of components and empirical correlations of the mean heat transfer coefficient of natural convection according to flow pattern of fluid. Concerning heat transfer coefficients of natural convection between the outer surface of the shell of the pressure vessel and the side cooling panel (Region 1 in Fig. 3.21) as well as between the side surface of the simulated core and the inner surface of the pressure vessel (Region 2), the mean Nusselt numbers based on characteristic length R are calculated by extrapolation of the empirical correlation by Thomas and de Vahl Davis [37] for inner and outer cooling concentric cylinders, as shown below: NuR 5 0:28Ra0:258  Pr0:006  K 0:442 ðRa , 2 3 105 ; 1 , K , 10; 1 , H , 33; 0:5 , Pr , 104 Þ:

(3.16)

This correlation is selected because the application ranges of K and H satisfy the present experimental conditions. The coefficients K and H are ratio between outer and inner radii of a concentric cylinder ( 5 rc/ri), and aspect ratio ( 5 L/(rori)), respectively. Hence, ri and ro are radius of the inner wall of a concentric cylinder (m) and radius of the outer wall of a concentric cylinder (m), respectively. Characteristic length R is the distance between the inner heating and outer cooling cylinders.

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High Temperature Gas-cooled Reactors

Regarding natural convective heat transfer coefficients for the top surface of the simulated core (Region 3), upper cooling panel (Region 4), and outer surface of the upper head of the pressure vessel (Region 5), the mean Nusselt numbers based on diameter d are calculated from the empirical correlations by Fishenden and Saunders [36] for upper heating and lower cooling horizontal surfaces as shown below:   Nud 5 0:54ðRad Þ1=4 105 , Rad , 2 3 107 ;   Nud 5 0:14ðRad Þ1=3 2 3 107 , Rad , 3 3 1010 :

(3.17)

Concerning natural convection heat transfer coefficients for the inner surface of the upper head of the pressure vessel (Region 6), the mean Nusselt numbers based on radius r of the upper head are calculated from the empirical correlations by Shiina et al. [39] as shown below:   Nud 5 0:1974ðRar Þ0:25 106 , Rar , 109 ;   Nud 5 0:0312ðRar Þ0:33 109 , Rar , 5:5 3 1010 :

(3.18)

The correlations are selected because the application ranges of the correlations cover the present experimental conditions. Regarding natural convective heat transfer coefficients for the bottom surface of the simulated core (Region 7), the inner surface of the lower head of the pressure vessel (Region 8), and the lower cooling panel (Region 9), the mean Nusselt numbers based on the diameter, d, are calculated from the empirical correlation by Fishenden and Saunders [38] for upper cooling and lower heating horizontal surfaces as shown below: Nud 5 0:27ðRad Þ1=4

(3.19)

Concerning natural convection on the outer heating surface of the lower ellipsoid head of the pressure vessel (Region 10), the mean Nusselt number based on longer radius r of the lower head of the pressure vessel is calculated from the empirical correlation by Theofanous [40] for the hemisphere heating surface as shown below: Nur 5 0:206  Rar 0:303  Pr0:084

(3.20)

Regarding natural convection heat transfer coefficients of the standpipes (Region 11) as vertical columns with lateral heating surfaces, the mean Nusselt number based on length x of the standpipe is calculated from the empirical correlation by Saunders [41] as shown below: Nux 5 0:59ðGrx  PrÞ1=4

(3.21)

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Table 3.4 Nusselt numbers of components used in numerical simulations [30]. Region number in Fig. 3.21

1 2 3 4 5 6 7 8 9 10 11

Cases 1 and 2

Case 3

Rayleigh number

Mean Nusselt number

Rayleigh number

Mean Nusselt number

4.35 3 108 5.65 3 106 9.43 3 107 2.12 3 1010 2.66 3 109 8.92 3 107 9.43 3 107 4.37 3 108 2.66 3 109 2.03 3 108 

38.0 11.4 63.7 387 193 19.2 26.6 39.0 61.3 65.6 

4.06 3 108 8.20 3 105 4.69 3 106 3.25 3 1010 7.80 3 108 1.08 3 107 4.70 3 107 1.18 3 108 5.45 3 109 1.54 3 108 4.77 3 107

37.4 6.93 23.4 443 129 11.3 22.4 28.2 68.0 60.4 46.5

The mean Nusselt and Rayleigh numbers employed in the present simulation are summarized for each region in Table 3.4. The simulation is carried out to evaluate the temperature distribution of the pressure vessel, based on the boundary conditions provided by the experimental heat flux data of the simulated core, and the measured temperature conditions of the cooling panels.

3.3.4 Evaluation of hot spot by natural convection Fig. 3.25 shows the experimental and numerical results of the temperatures of the upper head and the upper part of the shell of pressure vessel in case that helium gas is evacuated from the pressure vessel thus the natural convection inside the pressure vessel is restricted. Temperatures of the pressure vessel shell are nearly uniform with the exception for the flange at a distance of 1.0 m from the top of the upper head. Temperatures of the upper heads show almost uniform and lower than those of the shell. Temperatures on the cooling panels show almost uniform around 25 C that is close to the temperature of the cooling water. Numerical results of the temperatures of the THANPACST2 code well estimated the experimental ones on the upper head and the shell of the pressure vessel under the boundary condition of the experimental data for the heat flux of the simulated core and the temperatures of the cooling panels. The numerical results by the other safety evaluation codes MORECA and THERMIX also well estimated experimental results [33]. As far as natural convection flows only outside the pressure vessel, the three codes fairly well agreed with the experimental results of the temperatures on the pressure vessel.

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High Temperature Gas-cooled Reactors

Figure 3.25 Experimental and numerical results of temperature of pressure vessel under condition of vacuum in pressure vessel (Case 1) [30].

Figure 3.26 Experimental and numerical results of temperature of pressure vessel under condition of helium gas in pressure vessel (Case 2) [30].

Fig. 3.26 shows the experimental and numerical results of the temperatures of the upper head and the upper part of the shell of pressure vessel in case that helium gas is filled with the pressure vessel; thus, the natural convection of superheated gas affects the temperatures of the pressure vessel. Temperature profile of the cooling panel is similar to the case that helium gas is evacuated from the pressure vessel. In the experiment, the highest temperature that appears on the top of the upper head of the pressure vessel is 210 C and is 60 C higher than that in the vacuum condition in which the natural

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convection flow is restricted in the pressure vessel. This indicates that the natural convection flow of helium gas in the pressure vessel superheats the top of the upper head of the pressure vessel. Numerical results for temperatures of the upper head of pressure vessel by the THANPACST2 code well estimated the experimental results within the errors of 229 C under the boundary condition of the experimental data for the heat flux of the simulated core and the temperatures of the cooling panels. The ratio of thermal radiation heat transfer to the total heating rate is 60% on the simulated core. The numerical results of the safety evaluation codes MORECA and THERMIX, which did not calculate the local natural convective flow inside the pressure vessel, were significantly lower on the upper head of the pressure vessel in comparison with the experimental values [33]. Thus it was revealed that it is necessary to calculate the natural convective local flow of superheated gas flowing upward from the lower part inside the pressure vessel for precise predictions of the pressure vessel temperature. The temperature contour and velocity vector maps of the gases inside and outside of the pressure vessel acquired from the calculation by the code THANPACST2 are shown in Fig. 3.27. The numerical results indicate properly that helium gas superheated by the simulated core flows upward and superheats the top of the upper head of the pressure vessel where the highest temperature appears in the experiment, as shown in Fig. 3.27A. However, helium gas temperature decreases significantly and is 70 C lower than the experimental value near the point close to the top of the upper head of the pressure vessel. Helium gas superheated by the simulated core flowing upward is disturbed by the stepwise meshes of the upper head of the pressure vessel in the numerical model and flows downward before it reaches the top of the upper head of the pressure vessel, as shown in Fig. 3.27B. The heat transfer area of the upper head of the pressure vessel in the numerical model is 1.2 times larger than the actual value. As a result, temperatures at the upper head of the pressure vessel and in helium gas around the top of the upper head are lower than the experimental values. The air heated below the lower head of the pressure vessel flows through the airflow channels at the lower and upper edges of the simulated skirt-type support. It is superheated and flows along the shell of the pressure vessel. The air is cooled and flows down along the upper and the side cooling panels. The air forms a large circulating flow and the maximum velocities of the upward and the downward air flow are around 0.9 and 0.5 m/s, respectively.

3.3.5 Evaluation of local hot spot around standpipes Fig. 3.28 shows the experimental results of the temperatures of the pressure vessel. Temperatures of the pressure vessel appear lower below the elevation of 1.5 m at which heater segment No. 5, which is powered off, is located. Temperatures of the upper head increase within the distance of 0.5 m from the top and peak at 430 C on the top. The temperature of helium gas just below the top of the upper head is measured at 540 C. This confirms that natural convection of high temperature helium gas superheats the upper head of the pressure vessel. Thermal radiation heat transfer from the upper head to the upper cooling panel is restricted by the standpipes. This is confirmed by the fact that the temperature of the upper head increases rapidly up

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High Temperature Gas-cooled Reactors

Figure 3.27 Numerical results under the condition of helium gas in pressure vessel (Case 2) [30]. (A) Temperature contour. (B) Velocity vector.

to 430 C on the upper head, since the standpipes restrict thermal radiation heat transfer from the upper head to the upper cooling panel and at the same time reduce the heat transfer area on the upper head. The numerical results of the pressure vessel temperature by the THANPACST2 code are shown in Fig. 3.28. The numerical result shows a good agreement with the experimental result for the temperature profile on the upper head, which increases steeply within the distance of 0.5 m from the top. However, the conservative assumption employed in the numerical model, which sets the area occupied by the standpipes modeled with a thermal conductivity of thermal insulation on the upper head to be two times larger than the actual one, resulted in an estimation of significantly higher peak temperature at the top of the upper head. Meanwhile, the calculated temperatures of the pressure vessel fall within a range of 225 C to 170 C

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Figure 3.28 Experimental and numerical results of temperature of pressure vessel under condition of helium gas in pressure vessel with standpipes (Case 3) [30].

from the experimental values. The thermal radiation heat transfer accounts for an estimation of 81.8% in the total heat input from the simulated core. On the other hand, the safety evaluation code THERMIX did not show the steep increase of temperature in the standpipe region on the upper head. To evaluate the effect of natural convective air flowing through the standpipes on the temperature distribution of the pressure vessel, the numerical results of the THANPACST2 code using porous body cells and solid cells to simulate the standpipes were compared as shown in Fig. 3.28. The numerical result of the pressure vessel temperature using the solid cells to simulate the standpipes is 70 C higher than that using the porous body cells because the flow above the top of the pressure vessel is disturbed by the standpipes, which is simulated by the solid cells. Thus it was revealed that the natural convection of the air flowing through the standpipes cools the pressure vessel effectively.

3.4

Intermediate heat exchanger

3.4.1 Introduction The IHX of the HTTR is a 10 MW (thermal) vertical counter flow type heat exchanger to exchange heat between the primary and secondary helium gases. Fig. 3.29 shows an isometric view of the IHX. The IHX has 96 helically coiled tubes separated into 6 layer bundles in the radial direction, surrounding a center pipe. Each tube bundle is fixed as a hot header, which is located at the bottom end of the center pipe, and is supported by 6 tube supports in the tangential direction. Radiative plates are installed between each bundle to promote heat transfer by radiation. The center pipe and tube bundles move freely in the vertical and radial directions to accommodate any thermal displacement, but transmit lateral seismic loads to the shell through stoppers. The primary helium gas of 950 C and 4 MPa flows upward on the shell

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High Temperature Gas-cooled Reactors

Figure 3.29 Schematic view of intermediate heat exchanger [42].

side, and the secondary one of 300 C and 4.1 MPa flows downward in the helically coiled tubes and leaves through the hot header and center pipe [42]. Experimental and analytical studies were carried out to confirm the structural integrity of the IHX as follows: 1. 2. 3. 4. 5.

Creep collapse of the tube against external pressure, Creep fatigue of the tube against thermal stress, Seismic behavior of the tube bundles, Thermal hydraulic behavior of the tube bundles, and In-service inspection technology of the tube.

3.4.2 Creep collapse of the tube against external pressure 3.4.2.1 Objective and test procedure In the case that the secondary helium gas is lost in a depressurized accident, the helically coiled tubes coaled the possibly creep collapse by the external pressure of the primary helium gas, 4 MPa, due to the high temperature. Experimental and analytical studies on creep collapse of the helically coiled tubes against external pressure were carried out [43,44]. Material of the tube is a nickel-based superalloy, Hastelloy XR, which is a modified version of Hastelloy X to improve creep strength and corrosion resistance in the helium environment of the HTGR.

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The dimensions of the test tubes are 31.8 mm in outer diameter, 400 mm in length, and 24 mm in thickness. The ovality of the test tube, Φ, is defined by the following equation: Φ5

Rmax 2 Rmin 3 100%; R

(3.22)

where Rmax, Rmin, and R are the maximum and the minimum outer radii, and the mean outer radius (515.9 mm), respectively. The initial ovality of the test tubes was approximately 0%, 3%, or 6%. During experiments, the external pressure of 5 or 6 MPa was imposed on the test tubes at 950 C.

3.4.2.2 Test results Fig. 3.30 shows relationship between the collapse time and thickness of the tube. It was found that the collapse time became shorter as the initial ovality increased and the wall thickness decreased. The experimental results were compared with the analytical ones shown with lines in the figure. Constant stress creep tests for forged material were carried out by a balance weight inside an electric-heated furnace at 950 C in air. Next, a Norton-type creep constitutive of Hastelloy XR was established for the further analyses of creep deformation: ε_ c 5 3:3025 3 10211 σ4:8458 ;

(3.23)

where ε_ c and σ are creep strain rate (1/h) and applied stress (MPa). Young’s modulus at 950 C was taken to be 1.285 3 105 MPa, and the elastic and inelastic Poison’s ratio were taken to be 0.3 and 0.5, respectively.

Hastelloy XR 950 C

Collapse time (h)

103

Φi=0% Φi=6%

102 Open symbols : 5 MPa Solid symbols : 6 MPa

101

100

Φ i=0.1%, 5 MPa (ABAQUS) Φ i=0.1%, 6 MPa (ABAQUS) Φ i=6%, 5 MPa (ABAQUS)

2

3 Thickness (mm)

4

Figure 3.30 Relationship between collapse time and thickness of the tube [43].

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High Temperature Gas-cooled Reactors

Using Eq. (3.23), the creep collapse times were analyzed by the FEM code, ABAQUS. It was confirmed that the analyses by FEM code reasonably predicted the collapse time. In the after-inspection, many cracks on the outer surface of a few collapsed test tubes were observed by liquid penetration examination. Leak tests with air were carried out to confirm whether or not the crack passed through the tube wall. Leak was not observed for all the collapsed tubes. On the outer surface, the microcracks were mainly observed to have grown along grain boundary. The crack depth was less than 0.6 mm. It is concluded that the leak tightness can be maintained despite a collapse of the tube.

3.4.3 Creep fatigue of tube against thermal stress 3.4.3.1 Objective and test procedure A lower connecting tube having a complex configuration is welded between the hot header and the helically coiled tube. In normal operation, the lower connecting tube would be subjected to thermal stress caused by the difference of thermal expansion and temperature between the hot header and the helically coiled tube. Experimental and analytical studies on creep fatigue of the lower connecting tubes against thermal stress were carried out. A full-scale mock-up model is composed of the hot header, the lower connecting and helically coiled tubes, as shown in Fig. 3.31. The connecting tube is welded into the sub of the hot header, and the helically coiled tube was fixed to a support

Helically coiled tube

Support

Upper elbow Lower connecting tube Lower elbow

Hot header

Stub (mm)

Tube: 31.8O.D.x3.5t Hot header: 450O.D.x75t

Figure 3.31 Schematic view of full-scale mock-up model [42].

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Figure 3.32 Supporting positions of tubes [42].

with a load transducer. The model has eight lower connecting and helically coiled tubes separated into four types, namely AD; angles between the stub and the support are 90 , 120 , and 150 , respectively, as shown in Fig. 3.32. The center pipe was cyclically moved down and up in the vertical direction in the helium atmosphere of 950 C to simulate thermal expansion stress. The displacement between the hot header and the support, the displacement rate, and holding time at the maximum displacement were 50 mm, 120 mm/h, and 10 min, respectively, and the cycles of the displacement were more than 10,000. This test condition is severer than the actual one of the IHX.

3.4.3.2 Test results The tubes of the types A and D with the supporting angle of 90 degrees were fractured at the 4756th and 11,260th cycles, respectively. The other six tubes were not fractured; however, the lower connecting tubes were inclined and the helically coiled tubes hung down by empty weight. The crack in the type A tube was observed at the welding position between the stub and the lower connecting tube, and the crack with 15 mm length and 0.6 mm width was perpendicular to the axis of the tube. The type D tube was fractured at the parent material of the lower

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High Temperature Gas-cooled Reactors

elbow. The cracks occurred mainly along the grain boundary at the outer surface for both tubes. The elastic-creep analyses by the FEM code, ABAQUS, was performed to predict the life time of the tubes and compared with the experimental results. A Garofalo type creep constitutive equation was established as follows:   εc 5 εt 1 2 e2rt ε_ m t;

(3.24)

where εc , εt , r, t, and ε_ m are the creep strain, the maximum creep strain, the reciprocal of time constant, the time, and the minimum creep strain rate, respectively. Two type creep constitutive equations, that is, Eqs. (3.23) and (3.24), were applied for the analyses. The creep constitutive Eq. (3.22) was derived from the constant-creep tests and the creep constitutive Eq. (3.24) was used for the design of the IHX. The life time was predicted using the analytical results for the first three cycles in accordance with the linear damage rule [45] as follows: 3 X

  ðDc Þi 1 ðDc Þ3 3 Nf 2 3 5 1;

(3.25)

i51

where (Dc)i, (Dc)3, and Nf are the creep damage at the ith cycle, one at third cycle and the cycle to failure, respectively. The tube was modeled by three dimension 2-node elbow elements. Fig. 3.33 shows the comparison between the predicted and experimental life time. It was found that the safety factor of the predicted life time with the creep constitutive Eq. (3.23) and the linear damage rule was more than 10 times compared with the experimental. In case of the creep constitutive Eq. (3.24), the safety factor was more than 100. It was considered that the Mises equivalent stress history

Experimental life, Nf,exp (cycle)

105 104 103 Factor of 100

102 Factor of 10

101

Nf, pre=Nf,exp Creep constitutive equation Stress constant condition Design condition

100 100

101 102 103 104 Predicted life, Nf,pre (cycle)

105

Figure 3.33 Comparison between predicted and experimental life time [42].

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and the creep damage per cycle by the creep constitutive equation-2 were larger than those by the creep constitutive Eq. (3.23). Therefore it was found that the prediction of the life time by the elastic-creep analyses using the Garofalo type creep constitutive equation and the linear damage rule were very conservative compared with the experimental results in this test.

3.4.4 Seismic behavior of tube bundle 3.4.4.1 Objective and test procedure The seismic test was carried out using a partial model of the tube bundle to focus on investigating the vibration behavior induced in the IHX through an interaction between the tube bundle and the seismic stop [46]. Fig. 3.34 shows the structure of the model, which has a partial bundle divided into three layers of the tubes: an outer layer of 1.3 m diameter, a middle layer of 1.2 m, and an inner layer of 1.1 m. It contains 54 helically coiled tubes, supports, the center pipe, and lower connecting tubes. Additionally, for the IHX, the seismic stop is fixed on the shell; in the test, it is modeled by a steel rod and fixed on the structural steel diagonal bracing mounted on the shaking table. The gap between the tube support and the seismic stop was set at 2 mm, estimated as the largest gap in the IHX. The model was excited on the shaking table, having a size of 4.5 3 4.5 m, the maximum load force was 500 kN by using sine sweeps of 0.12 and 0.25 m/s2. The

Figure 3.34 Structure of seismic test model [46].

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sweeps were low enough to achieve a steady condition and an equivalent ground motion induced by the S1 seismic wave [47], that is, the design of the estimated strongest earthquake for the HTTR.

3.4.4.2 Test results The acceleration response spectra on the bottom end of the center pipe are shown in Fig. 3.35 together with the analytical results described later. The frequency of jump appeared at 4.9 Hz for the input acceleration of 0.12 m/s2 and at 5.4 Hz for an acceleration of 0.25 m/s2. The larger the input acceleration was, the higher the frequency of the jump phenomena. As for the acceleration response spectra of the inner and the outer layers of the tubes, it can be said that each layer of tubes vibrated together as one bundle, and the response curve could be identified with that of the center pipe. Accordingly, the center pipe and the bundle may be treatable as a completely stiff system. The response spectra associated with the time history of the impact on the seismic stop in sine sweeps of 0.12 and 0.25 m/s2 are shown in Fig. 3.36. The

Figure 3.35 Acceleration response spectra of center pipe [46].

Impact force (N)

Center pipe 104

103

Position

Input acc. (m/s 2)

Right

0.12 0.25

Left

0.12 0.25

Analysis 0.12 0.25

Gap 2 mm

102 100

101 Frequency (Hz)

Figure 3.36 Response spectra of impact force on seismic stop [46].

4x101

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maximum value was generated at the resonance frequency of the center pipe with the bundle. To estimate the impact force on the seismic stop, the analyses were carried out using a simple model in which the center pipe and the bundle were treated as a one-dimensional spring mass since the experimental results confirmed that the tube bundle vibrated together with the center pipe. Then, the motion of the center pipe is given by the following considerations: If; ..jX1 2 X2 # G;j Mc X1 5 Kc X1 1 Cc X_ 1 ;

(3.26)

or if jX1 2 X2 j . G, .. Mc X1 5 Kc X1 1 Kb ½X1 2 X2 2 signðX1 2 X2 ÞG 1 Cc X_ 1 ;

(3.27)

Kb ½X1 2 X2 2 signðX1 Þ 2 ðX2 ÞG 5 Ks X2 ; signðX1 2 X2 Þ 5 1; for X1 2 X2 . 0; and signðX1 2 X2 Þ 5 2 1; for X1 2 X2 # 0 where Mc, Kc, Cc, Kb, Ks, G, X1, and X2 are the equivalent mass of the center pipe including the tube bundle, the equivalent stiffness, the damping factor, the tangential stiffness of the bundle, the stiffness of the seismic stop, the size of the gap between the seismic stop and the bundle, the displacement of the center pipe including the bundle and the displacement of the ground motion, that is, the shaking table or the shell, respectively. Assuming the center pipe oscillates as a pendulum together with the bundle, the equivalent mass Mc is given by Mc 5

I ; L2

(3.28)

where I and L are the inertia moment of the center pipe including the bundle with respect to the pivot at which the top end of the center pipe connects with the frame, and distance between the pivot and the contact position with the seismic stop, respectively. Fig. 3.36 shows comparisons of the analytical results with the experimental ones concerning the acceleration response spectra on the center pipe and the impact force on the seismic stop. The analytical results, particularly at the spectra peak frequencies, gave excellent agreement with the experimental ones despite the amplitude of input acceleration. As for the frequency responses, the discrepancy between the analytical and experimental results in the range from ,3 Hz or .6 Hz is considered to result from two-dimensional vibrational components associated with rotation of the

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center pipe, deviation of the mass, etc., which cannot be represented by the simple one-dimensional model. The simple model, however, can be said to be suitable for the estimation of the resonance characteristics of the IHX with the bundle tubes.

3.4.5 Thermal hydraulic behavior of tube bundle 3.4.5.1 Objective and test procedure An experimental study was carried out to investigate the flow-induced vibration of the tube bundle due to the flow velocity of the tube outside for demonstration of the structural integrity of the tube bundle, and heat transfer characteristics on forced convection of the tube outside and radiation by radiative plates [48]. The test was conducted using the same full-scale partial model and air instead of helium gas. The test model, as shown in Fig. 3.37, has a center pipe, a hot header, and 54 helically coiled tubes, which are separated into three layer bundles. The radiative plates are installed between each bundle to promote heat transfer.

Figure 3.37 Structure of thermal hydraulic test model [48].

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Flow rates of the air, temperatures of the metal and air, and vibration behavior of the tube bundles were measured with vortex flow meters, thermocouples, and accelerometers, respectively. The percentages of flow rate were 27.8%, 30.0%, and 42.2% for the primary air from the inner to outer bundles, and 31.5%, 33.3%, and 35.2% for the secondary air, respectively, according to the calculation based on the flow areas. The Reynolds number, Re, of the primary air is 6.0 3 1032.66 3 104 (IHX: 1.54 3 104) based on the outer diameter of the tube and the air velocity in minimum gap of the tube outside, and the Reynolds number of the secondary air is 4.7 3 104 (IHX: 5.02 3 104) based on the inner diameter.

3.4.5.2 Test results

Acceleration,G (m2/s)

The vibration behavior of the center pipe and three tube bundles in air flowing at a room temperature with the primary air Reynolds number of 1.54 3 104 is shown in Fig. 3.38. The tube bundles vibrated together with the center pipe and individually vibrated at frequencies, which corresponded to each natural frequency. The vibration was amplified at constant frequencies by increasing the primary air velocity. The vibration amplitude was calculated from the obtained response acceleration and frequency. The maximum dimensionless amplitude of the tube bundle, y/do (y: vibration amplitude, do: outer diameter of the tube) was as small as B2.5 3 1023. According to the study on vortex shedding of in-line tube arrangement by Chen [49], there is a possibility that the vortex shedding frequency coincides with the natural frequency of the tube bundle, that is, a “lock-in.” The fluidelastic instability was also investigated according to studies by Connors [50], Pettigrew et al. [51], Paidoussis [52], and so on. The test data stayed in the stable region provided by Pettigrew et al. On the other hand, there was a possibility of the fluidelastic instability according to Paidoussis. No abnormal vibration, however, was observed in the test conditions although the vibration increased as the flow velocity increased. Therefore it is concluded that no problem exists for the fluidelastic vibration of the tube bundles in the IHX. The tubes receive heat from the primary air by forced convection and the radiative plates by radiation, and transfers the heat to the secondary air. Fig. 3.39 shows Natural frequency of tube bundle Natural frequency of center pipe 0.4

Outer bundle Middle bundle Inner bundle Center bundle

0.0 0

50 Frequency, f (Hz)

100

Figure 3.38 Vibration behavior of tube bundles and center pipe due to primary air flow (Reo 5 1.54 3 104) [48].

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Figure 3.39 Heat transfer characteristics of tube outside by forced convection [48].

the obtained heat transfer characteristics outside the tubes by forced convection at inlet temperatures of the primary air from 200 C to 300 C, compared with the heat transfer coefficients of the crossflow around the tubes for the in-line arrangement proposed by Zukauskas [53] and Fishenden and Saunders [38]. The obtained heat transfer coefficients were 6%20% smaller than those by Zukauskas and FishendenSaunders, because of the tube incline in a flow direction of 12 degrees, and can be given by 0:3 Nuc 5 0:78Re0:51 o Pr ;

(3.29)

where Nuc, Reo, and Pr are the Nusselt number by forced convection, the Reynolds number based on the outer diameter of the tube, and the Prandtl number, respectively. The radiative plates provide us not only with promotion of the heat transfer but also increase in the pressure drop. In order to estimate the advantage of the radiative plates, the relationship between the Nusselt number and the pumping power needs to be obtained. As a result, it can be said that the radiative plates have the merit of promoting the heat transfer despite the disadvantage of increasing the pressure drop.

3.4.6 In-service inspection technology of tube 3.4.6.1 Objective and test procedure The structural integrity of the IHX tubes is very important for the safety of the HTTR because they retain the primary pressure boundary. An in-service inspection

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Figure 3.40 Cross-section of the ECT probe (dimensions: millimeter) [54].

(ISI) of the tubes is carried out with eddy current testing (ECT). A performance test of the ECT probe for discontinuities in the base metal and welded joints was conducted [54]. The ECT probe has a differential and self-induction type of coil arrangement because it is required to travel in helically coiled tubes, and this design reduces the influence of the curvature of the tubes. The probe has a 20-mm outer diameter, is 40 mm long, and has two 16 mm-O.D. coils as shown in Fig. 3.40. The Hastelloy XR test tubes including base metal and welded joints have the same dimensions as the IHX tubes. Artificial discontinuities of the test tubes were made with reference to the ASME standard. A 1.7-mm diameter 100% through-wall hole and 1.5-mmwide 360 degree circumferential grooves .10% through from the inner and outer tube surfaces were selected as the standard discontinuities. Moreover, the inspection performance for 1.7- and 3.2-mm diameter flat-bottom holes, a 0.5-mm diameter 100% through-wall hole, and 0.5- and 1.7-mm wide 90 degree circumferential grooves were investigated as shown in Fig. 3.41.

3.4.6.2 Test results Fig. 3.42 shows the relationship between the actual depths (δ) of discontinuities and predicted ones (δp) in the base metal obtained from signal of the ECT probe at frequency of 48 kHz. The predicted depths were different from the actual ones within 618%. There were no clear differences between the phase angles of the discontinuities in the straight and in the curved tubes. The ECT probe could detect small discontinuities such as the 0.5-mm diameter 100% through-wall hole, and 0.5- and 1.7-mm wide 90 degree circumferential grooves. As for the welded joints, the discontinuities were inspected at multiple frequencies of 48 and 180 kHz to remove the noise of the back-excess weld metal of the welded joints. The inspection performance was lower compared with the base metal, and the following small discontinuities were undetected: 1.7-mm diameter flat-bottom holes and 40% through 60% depth at the inner and the outer surfaces, 1.5-mm wide 90 degree grooves and 20% through 40% depth at the outer surface. The predicted depths were different from the actual ones within 620% of the wall thickness.

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High Temperature Gas-cooled Reactors

Figure 3.41 Test tubes including artificial discontinuities: (A) base metal and (B) welded joint (dimensions: millimeter) [54].

3.5

Basic feature of air ingress during primary pipe rupture accident

3.5.1 Introduction The inherent properties of HTGR facilitate the design of HTGR with high degree of passive safe performances, compared to other type of reactors. However, it is still not clear if the present HTTR type reactor can maintain a passive safe function during the primary-pipe rupture accident, or what would be a design criterion to guarantee HTGR with the high degree of passive safe performances during the accident. The primary-pipe rupture accident is one of the most common accidents related to the basic design regarding an HTGR, which has a potential to cause the destruction of the reactor core by oxidizing in-core graphite structures and to release fission products by oxidizing graphite fuel elements. It is a guillotine-type rupture of the double coaxial pipe at the nozzle part connecting to the bottom of the reactor pressure vessel, which is a peculiar accident for the HTTR type reactor. When the

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Figure 3.42 Comparison between actual and predicted depths of discontinuities in base metal at 48-kHz frequency [54].

primary-pipe rupture accident happens, one may assume that air entering into the reactor pressure vessel reacts with high temperature graphite structures and causes temperature rise of the reactor core and corrosion of the graphite components. Therefore it is very important to make sure that the air ingress process during the primary-pipe rupture accident cannot seriously oxidize the graphite fuel elements to release the radioactive materials from the reactor core to the environment nor severely damage the graphite components to lose the integrity of the reactor internals. A schematic drawing of the HTTR and the coolant passages in the reactor is provided in Fig. 3.43A. A hot leg consists of an inner passage of a coaxial duct, a high temperature outlet duct, a high temperature plenum, and fuel cooling channels. A cold leg consists of an annular passage of the coaxial duct, a bottom cover, and an annular passage between the reactor pressure vessel and permanent reflector. As the hot and cold legs are connected at the top space, they make a kind of reverse U-shaped tube as shown in Fig. 3.43B. When the primary-pipe rupture accident occurs, the high-pressure helium gas coolant in the reactor is forced out into the reactor containment through the breach. Gas pressure should become balanced between the inside and outside of the reactor pressure vessel. During this stage, called the depressurization stage, air is not able to enter the reactor core from the breach. After the depressurization stage, it is supposed that air enters the reactor core from the breach due to molecular diffusion and natural circulation of a multicomponent gas mixture induced by the distribution of gas temperature and the resulting concentrations in the reactor. Carbon monoxide (CO) and dioxide (CO2) are produced in the reactor because the oxygen (O2) contained in air reacts with the high temperature graphite structures. The density of the

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Reactor container Reactor pressure vessel Top space Replaceable reflector Core Permanent reflector High temperature plenum High temperature outlet duct Bottom space Coaxial duct (primary-pipe) Hot-leg

Cold-leg

(B) (A)

Flow direction at the primary-pipe rupture accident Flow direction at the normal operation

Figure 3.43 Schematic drawing of HTTR and model of coolant passages in reactor [55]. (A) HTTR. (B) Reverse U-shaped model.

gas mixture in the reactor gradually increases as air enters by the molecular diffusion and natural circulation of the gas mixture in the first stage of the accident. Finally, the second stage of the accident starts after the natural circulation of air occurs suddenly throughout the entire reactor [55]. Previous studies focused mostly on the mixing process of two-component gases by molecular diffusion and natural circulation in a reverse U-shaped tube and in a simple test model of the HTTR [55]. In order to investigate the basic features of the flow behavior of multicomponent gas mixture consisting of helium (He), nitrogen (N2), O2, CO2, CO, etc., experimental and numerical analysis studies were performed [56]. These studies are to clarify a combined phenomenon in terms of molecular diffusion and natural circulation of the multicomponent gas mixture along with a graphite oxidation reaction in a reverse U-shaped tube. The numerical analysis result was in good agreement with the experimental result in regard to the density change of the gas mixture, the molar fraction change of the gas species, and the onset time of the natural circulation of air. To understand the air ingress process in the HTTR and to validate the computer code, however, it is absolutely necessary to investigate the phenomena in a test facility that simulates the HTTR design. For these purposes, a test apparatus was constructed, which simulates the HTTR, and experiments simulating the primary-

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pipe rupture accident of the HTTR were carried out [57]. Based on the results of these fundamental studies, computer codes were developed to predict air ingress processes for various designs of the HTTR type reactor [58]. In order to predict or analyze the process of air ingress during the accident, it is very important to develop computer programs and to validate them by experiments. Furthermore, the objectives of this study are to investigate the air ingress process and to develop the passive safe technology for the prevention of air ingress. This section describes a computer program, developed to analyze the process of air ingress during the first stage of a primary-pipe rupture accident.

3.5.2 Basic feature of air ingress phenomena in a reverse Ushaped channel 3.5.2.1 Experimental apparatus, method, and results Experimental and numerical studies were performed on the combined phenomena of molecular diffusion and natural circulation in a two-component gas system (N2-He) in a reverse U-shaped tube. Fig. 3.44 shows an example of the mole fraction and the ingress velocity changes of N2 in the reverse U-shaped tube. The solid line in the figure is the ingress velocity obtained by the hot-wire anemometer attached at the lower end of the high temperature side pipe. The symbols (o) and (Δ), respectively, represent the mole fraction changes at the same distance (600 mm) from the lower ends of the both side pipes. The mole fraction of N2 at the high temperature side pipe is higher than that at the low-temperature side pipe. 1.Densityof the gas mixture just after the valve open 'ρ=0.063

Low temperature He

First stage (at 120 min) 'ρ=0.056

High temperature He

ρ=0.156

2.Molecular diffusion and natural circulation (first stage)

ρ=0.093

Molecular diffusion

Low temperature N2

N2 Onset of ordinary natural circulation

Ingress velocity (m/s)

Heated

Cooled

Mole fraction of N2

Low temperature N2

Second stage

348°C

He

1

ρ=0.494

ρ=0.550

ρ=1.183 First stage (165 min)

High temperature gas mixture

Low temperature gas mixture

0.2

3.Ordinary natural circulation of N2 (second stage) Second stage (at 170 min) 'ρ=0.441 Low temperature N2

High temperature N2

ρ=1.085

ρ=0.644

0 0

60

120

180

Low temperature N2

Elapsed time after the valve open (min)

Figure 3.44 Basic feature of nitrogen ingress process in a reverse U-shaped tube [56].

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High Temperature Gas-cooled Reactors

This is because N2 is transported upward by the molecular diffusion from the both ends of the pipes and by the natural circulation of the gas mixture having the very slow velocity from the high temperature side to the low-temperature side pipe. According to the results obtained, the density of the gas mixture in the reverse U-shaped tube gradually increases as N2 enters as a result of molecular diffusion and because of a very weak natural circulation of the gas mixture. The calculated velocity of this very weak natural circulation of the gas mixture is about 10261023 m/s (1024 , Re , 1). The ordinary natural circulation of N2, which is about 0.2 m/s (Re . 200) in Fig. 3.44, occurs suddenly throughout the reverse U-shaped tube because the buoyancy force has risen to such an extent as to bring about the natural circulation. The time required from the first stage to the second stage is about 10 s in the ingress process. Therefore the production process of the second stage is a rapid phenomenon because the duration time of the first stage is about 165 min in the reverse U-shaped tube [55]. The helium injection method used this peculiar phenomenon is described in this section. It is one of some useful methods for prevention of air ingress.

3.5.2.2 Existence of parallel channels in a reactor core In the paragraph 3.5.2.1, the ingress process of N2 when He was filled in the reverse U-shaped tube composed of the high temperature and the low-temperature passages was described. However, it is necessary to investigate the influences of the local natural circulation in the parallel channels onto the process of N2 ingress during the primarypipe rupture accident because a lot of parallel channels with different temperatures exist in the reactor core. In this chapter, the influences of the local natural circulation in the vertical parallel channels onto the N2 ingress process are discussed [59]. The duration of the first stage of the N2 ingress process is also discussed from the experimental results of the reverse U-shaped passage, which has the parallel channels. An experimental apparatus is shown in Fig. 3.45. The apparatus consists of the reverse U-shaped passage having three vertical parallel pipes as the hot leg. The parallel pipes were heated and the other pipe was cooled. The lower ends of the heated side and cooled side channels were connected to a storage tank through a cutoff valve. The upper and lower ends of the parallel pipes are connected with a cylindrical plenum. The inner diameter of all pipes is 52.7 mm. The height of the heated section is 724 mm. The inner diameter of the plenum is 288 mm and height is 46 mm. The storage tank is a cylindrical container of 1000 mm in the inner diameter and 400 mm in height. The copper tube for cooling water was inserted into the container to keep the constant temperature of the gas. Fig. 3.46 shows the measurement points of temperature and density of the gas mixture. The mole fraction of N2 was obtained by measuring the speed of sound of the gas mixture. In addition to that, to obtain the mole fraction of N2, the density of gas mixture was measured with a density meter. The principle of measurement of this density meter is that the resonance frequency of the thin cylinder changes according to the density of gas mixture in cylinder surroundings. To measure the ingress flow velocity of N2, the hot-wire anemometer was inserted into the tube entrance.

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Figure 3.45 Experimental apparatus [59].

The experiment procedure is as follows: The cutoff valves installed between the reverse U-shaped passages and the storage tank were shut. Helium was filled in the passages and N2 was filled in the storage tank. Next, the wall of the parallel channels was heated until the temperature of the wall and the gas reached the test condition. Then the temperature was kept constant afterward. Pressure in the passages was kept at atmospheric pressure during heating. The cutoff valves were

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High Temperature Gas-cooled Reactors

Figure 3.46 Position of temperature and gas concentration measurement [59].

opened at the same time when the temperature of the wall and the gas reached the steady state. As for experimental conditions, four cases were set by using the temperature of the heater wall as a parameter. Table 3.5 shows the average temperatures of each part in the experiment. The wall temperature of the parallel channels and the gas temperature of the horizontal pipe are the integrated average temperatures of the axial direction. The others are the arithmetic mean temperatures.

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Table 3.5 Average temperature in typical experiments [59]. Condition

Isothermal/C1

Different/C2

Different/C3

Different/C4

CH-1CH-2CH-3 Duration CH-1 wall CH-2 wall CH-3 wall Lower plenum gas Upper plenum gas Horizontal gas Storage tank gas

281279280 3 h 42 m 281.2 279.4 280.1 139.2 183.0 213.4 23.4

281280211 2 h 54 m 281.3 280.1 211.4 150.3 212.7 213.3 23.5

246319245 2 h 45 m 245.8 318.7 245.3 146.7 207.3 183.0 22.0

284181213 2 h 54 m 283.5 180.6 212.9 128.6 201.5 122.0 19.1

Figure 3.47 Mole fraction change of nitrogen (C1) [59].

3.5.2.2.1 Isothermal condition Fig. 3.47 shows the mole fraction change of N2. The average temperature of each region is shown in Table 3.5 as the condition 1. The circle (o) symbol in the figure denotes the mole fraction of N2 of the measuring point 4 in the channel 1. The triangle (Δ) and rectangle (&) symbols indicate the mole fraction of N2 of the measuring point 7 in the channel 2 and that of the measuring point 10 in the channel 3, respectively. They are located 830 mm from the storage tank. The diamond (◄) and cross ( 3 ) symbols indicate the mole fraction of N2 of the measuring point 13, which is located 1130 mm from the storage tank and that of the measuring point 19, which is located 460 mm from the storage tank of the cooled channel. Though the distance from the storage tank is the shortest, the mole fraction of N2 of the point 19 is lower than that of the points 4, 7, and 10 at 120 min. Finally, the mole fraction of N2 of the point 19 is lower than that of the point 13, which is located the most far from the storage tank. This is because that N2 will be transported by the very weak natural circulation and molecular diffusion as same as the

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High Temperature Gas-cooled Reactors

case of the reverse U-shaped tube. Then, the amount of transported N2 by the natural circulation will increase with increasing time. The direction of the natural circulation flow is same as that of the molecular diffusion in the high temperature side. In the lowtemperature side, however, the flow is the counter direction of the molecular diffusion. Therefore the mole fraction of N2 not so much increases. In the case when the average temperature of the parallel channels is 280 C, the natural circulation of N2 occurs after 222 min. Then, the ingress process shifts to the second stage. Fig. 3.48A and B shows the temperature distributions of the heated wall and gas mixture in the channel 1. The localized natural circulation of the gas mixture will not occur in the parallel channels because the wall temperature will be same as the gas temperature. As shown in Fig. 3.48B, the gas temperature at the channel inlet decreases after 224 min because the natural circulation will occur in the apparatus.

Figure 3.48 (A) Temperature distribution of the heated wall (C1) [59]. (B) Temperature distribution of the gas mixture (C1) [59].

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Density difference between the heated and the cooled channel (kg/m3)

0.45 0.4 0.35 0.3 0.25 0.2 Average temp.of the pipe wall

0.15

172.7 (°C) 212.7 (°C)

0.1

258.6 8 (°C) 299.5 (°C)

0.05 0

346.6 (°C)

0

30 60 90 120 150 180 210 240 270 300 330 360 390 Elapsed time after the valve is opened (min)

Figure 3.49 Relationship between the duration of the first stage and the density difference [59].

Figure 3.50 Mole fraction change of nitrogen (C2) [59].

Fig. 3.49 shows the relationship between the duration of the first stage and the density difference of the heated and cooled channel. The density difference between the gas mixture density of the heated channel and that of the cooled channel does not exceed about 0.26 kg/m3 in all the experiments. Therefore the density difference of the gas mixture suddenly was increased in just before the onset of the natural circulation.

3.5.2.2.2 Different temperature condition Fig. 3.50 shows the mole fraction change of N2 in the different temperature condition. As shown in the figure, the mole fraction of N2 of the parallel channels almost equal that of the upper part of the channels. This is because the localized natural

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High Temperature Gas-cooled Reactors

circulation of the mixed gas will occur between the parallel channels in the different temperature conditions. Therefore the mole fraction of N2 of the upper plenum almost equals that of the lower plenum. Fig. 3.51A, B and 3.52A, B show the temperature distributions of the heated wall and gas mixture in the channels 1 and 3, respectively. The gas temperature at the lower end of the channel 1 is lower than the wall temperature as compared with the result of the isothermal condition. On the other hand, the gas temperature at the upper end of the channel 3 is lower than the wall temperature. Therefore the localized natural circulation of which the flow direction of the channels 1 and 2 is upward and that of the channel 3 is downward will occur in the parallel channels.

Figure 3.51 (A) Temperature distribution of the CH-1 heated wall (C2) [57]. (B) Temperature distribution of the CH-1 gas mixture (C2) [59].

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Figure 3.52 (A) Temperature distribution of the CH-3 heated wall (C2) [59]. (B) Temperature distribution of the CH-3 gas mixture (C2) [59].

The flow velocity of the localized natural circulation in the parallel channels will be faster than that of the very weak natural circulation in the reverse U-shaped tube; therefore the gas temperature in the lower and upper plena becomes equal. The duration of the first stage of the N2 ingress process was discussed from the results of the isothermal and nonisothermal experiments. Though the wall temperature of the channel 3 decreased at about 70 C, the duration of the first stage is shorter in the nonisothermal condition. The onset time of the natural circulation of N2 is about 174 min. Therefore the difference of the integration value of the density of the gas mixture between the high temperature channel and the low-temperature channel becomes large early when the natural circulation of the gas mixture is generated between the parallel channels. Then, the duration of the first stage is shorter and the ingress process shifts to the second stage.

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High Temperature Gas-cooled Reactors

Figure 3.53 Mole fraction change of nitrogen (C4) [59].

Figure 3.54 Relationship between average temperature of parallel channels and the duration of first stage [59].

Fig. 3.53 shows the experimental results of which the wall temperature of the parallel channel is kept at different temperature. The average temperature of each region is shown as Condition 4 in Table 3.5. In this case, the mole fraction changes of the parallel channels are almost same because the localized natural circulation generates in the parallel channels. Therefore the mole fraction of N2 in the channels and plenum almost equals. Fig. 3.54 shows the relationship between the average temperature of the parallel channels and the duration of the first stage. The ( ) symbol shows the duration in the isothermal condition (Condition 1). The (Δ, &, ◄) symbols show the duration G

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in the different temperature condition (Conditions 2, 3, 4). The temperatures written in the figure show the wall temperature of channels 1, 2, 3, respectively. As shown in the figure, the duration of the first stage is the longest when the temperature of the parallel channels is equal. The duration of the first stage becomes short when the localized natural circulation is generated between the parallel channels. Therefore it was found that the duration of the first stage was affected not only by the average temperature of the entire channels but also strongly by the localized natural circulation in the parallel channels.

3.5.3 Basic feature of air ingress phenomena in a simulated reactor apparatus 3.5.3.1 Numerical analysis FLOWGR provides a numerical method for analyzing the transient thermal hydraulic behavior by solving the one-dimensional basic equations for continuity, momentum conservation, energy conservation of the gas mixture, and the mass conservation of each gas species [58]. The main features of FLOWGR are explained in the following: 1. FLOWGR treats the gas mixture consisting of arbitrary components and takes into account the graphite oxidation and CO combustion reactions. The source and sink terms in the mass conservation equations resulting from these chemical reactions are expressed by the temperature-dependent Arrhenius formula having rate constants, which are specified by the user. 2. The reactor system to be analyzed by FLOWGR is modeled as a network of one-dimensional stream tubes, which represent components of the system such as a piping, plenum, and reactor core. Each stream tube is further divided into fluid volumes (control volumes or cells). There are three types of tube ends: normal, break, and closed end. The normal end connects with adjacent stream tubes, the closed end has no adjacent stream tube, and the break end represents the breach connected to the open air.

The basic equations (mass-based formulation) are as follows. The equation of continuity for the gas mixture: n @ρ @ðρuÞ X 1 5 Qi : @t @x i51

(3.30)

The equation of mass conservation for each gas species:   @ðρωi Þ @ðρuωi Þ @ @ωi 1 5 ρDi2m 1 Qi ; @t @x @x @x for i 5 1 to n 2 1, and ωn 5 1 2

Pn21 k51

ωk .

(3.31)

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High Temperature Gas-cooled Reactors

The equation of energy conservation:       @ ρcp T @ ρucp T @ @T Lh 1 5 λ 1 α ðTw 2 T Þ: @t @x @x @x Ae

(3.32)

The equation of momentum conservation:     @ðρuÞ @ ρu2 @p 1 f 1 2 ρg cos θ 2 ρujuj 52 : @t @x 2 De @x

(3.33)

The equation of state for the gas: pMm 5 ρRT

or

p 5 CRT:

(3.34)

Here, the friction factor (f) and the heat transfer coefficient (α) corresponding to the fully developed laminar flow are used [61]. The viscosity (μ) and thermal conductivity (λ) of each gas species and the gas mixture are obtained using the Wilke method and by the Eucken correlation [62], respectively. The other set of the basic equations (molecule-based formulation) includes the equations of continuity (Eq. 3.35) and of number of molecules conservation for each gas species (Eq. 3.36) instead of Eqs. (3.30) and (3.31). The equation of continuity for the gas mixture: n @C @ðCuÞ X Qi 1 5 : @t @x Mi i51

(3.35)

The equation of number of molecules conservation for each gas species:   @ðCXi Þ @ðCuXi Þ @ @Xi Qi 1 5 CDi2m : 1 @t @x @x @x Mi

(3.36)

P For i 5 1 to n 2 1, and Xn 5 1 2 n21 k51 Xk where Xi 5 mole fraction for the component gas i, Mi 5 molecular weight. In this set of equations, the conservation equation of momentum (Eq. 3.33) and the equation of state for the gas (Eq. 3.34) are used. However, the equation of energy conservation (Eq. 3.32) is omitted based on the assumption that the gas temperature is equal to the wall temperature. The diffusion coefficients Di2m for the multicomponent gas system are obtained from the diffusion coefficients for the binary gas system and the mole fractions of each gas species. 1 2 Xi j51 Xj =Di2j

Di2m 5 Pn

(3.37)

R&D on components

245

This coefficient is called the effective diffusion coefficient in the multicomponent gas mixture, which is provided by references [6064]. In the present analysis, the graphite oxidation reaction (CO2 reaction) and the carbon monoxide combustion (COO2 reaction) are taken into account. The dissipation or generation terms Qi of the mass conservation equations for O2, CO, and CO2 are written as: ð2Þ Qi 5 Qð1Þ i 1 Qi ;

(3.38)

ð2Þ where Qð1Þ i is the term for the graphite oxidation reaction and Qi is the one for the CO combustion. Using the Arrhenius formula, these terms are

  E1 ð1Þ Qð1Þ 5 K exp 2 ; i i RT Qð2Þ i

5 Kið2Þ exp



 E2 2 ; RT

(3.39)

(3.40)

where Kið1Þ and Kið2Þ are constants, and E1 and E2 are the activation energies. Details of these constants and activation energies are explained in reference [58]. The numerical methods are based on replacing the system of differential equations with a system of finite-difference equations fully implicit in time except the convection term for the momentum equation, which is treated explicitly. These results in a set of discretized nonlinear equations are solved by iteration procedure. In each iteration time, the implicit terms are formulated to be linear in the dependent variables at new time. The resulting linear systems are solved by direct inversion or a GaussSeidel iteration. The finite-difference equations used in FLOWGR can be obtained by integrating the conservation equations over a fixed fluid volume. The scalar properties (pressure, temperature, etc.) of the flow are defined at the volume centers, while vector quantity (velocity) is defined at the junctions. The equations of continuity and mass conservation (Eqs. 3.30 and 3.31), and the equation of energy conservation (Eq. 3.32) for the mass-based formulation as well as the equations (Eqs. 3.35 and 3.36) for the molar-based formulation are integrated with respect to the spatial variable x from the junction at xj to the junction at xj11. The equation of momentum conservation (Eq. 3.33) is integrated with respect to the spatial variable from the volume center at xj11/2 to the adjacent volume center at xj13/2. In this integration, loss of pressure is taken into account at the volume boundary with abrupt area change. The fluxes due to convection and diffusion, which appear in the integrated equations, are approximated using the hybrid difference scheme [65,66]. The finite-difference equations can be solved under the initial and boundary conditions. The initial conditions for the temperature, pressure, and mole fractions for each component gas are specified at the volume centers and those for velocities at the junctions. In order to analyze the phenomenon just after the depressurization stage, it is assumed that the gas velocities are zero for the initial condition. As for

246

High Temperature Gas-cooled Reactors

the boundary conditions, the pressure, temperature, and mole fractions of each gas species are specified at the end of the stream tube, which has the break end. The wall temperature of the stream tubes is also specified for each control volume as a function of time. At the closed ends of the stream tubes, the boundary conditions are that the flow velocity and fluxes due to convection and diffusion are zero. In order to obtain stable and fast convergence of the nonlinear equations, the equations of continuity and of momentum conservation for the gas mixture linearized by the pressure and velocity are solved using the SIMPLER method proposed by Patankar [65] along with the other conservation equations linearized by the mass (mole) fractions and temperature. In this iteration procedure, the numerical relaxation is applied for the production and dissipation terms as well as for the dependent variables in such a way. Γnew 5 δΓcal 1 ð1 2 δÞΓold

(3.41)

where Γ new is a new value, Γ cal is a calculated value without the relaxation, Γ old is an old value, and δ is a relaxation parameter, the value of which is chosen in a range 0 , δ , 1.

3.5.3.2 Comparison with experiments Fig. 3.55 shows an experimental apparatus, which consists of a reactor core simulator, a high temperature plenum, a water-cooled jacket corresponding to the reactor pressure vessel, and simulated inlet and outlet pipes corresponding to the coaxial tube. The reactor core simulator has four graphite pipes, a ceramic plenum, and electric heaters. The graphite pipe has an inner diameter of 40 mm and a length of 800 mm. The reactor core simulator divides into 12 temperature regions, which is the upper, middle, and lower parts of each graphite pipe. The graphite pipe temperature in each region can automatically be controlled individually. The inlet pipe simulates the outer pipe of the double coaxial primary pipe of the HTTR. It has an inner diameter of 69.3 mm. The outlet pipe simulates the inner pipe of the hightemperature primary pipe. An upper part of the outlet pipe is made of ceramic and a rest lower part is made of stainless steel. The inner diameter of the outlet pipe is 69.3 mm. At the bottom ends of the outlet and inlet pipe, ball valves are equipped to simulate the primary-pipe rupture accident. The pressure vessel simulates the reactor pressure vessel of the HTTR and it has a height of 2003 mm and an inside diameter of 920 mm. A partition plate is installed in the top space of the pressure vessel to lessen the volume of the top space, which shortens experiment time required for one run. The annular passage between the inner barrel and the pressure vessel has a gap of 60 mm. The main specifications of the experimental apparatus are provided in Table 3.6. In order to measure the ingress velocity of air, an ultra-sonic anemometer is equipped with the lower end of the outlet pipe. The temperatures of the gas mixture,

R&D on components

Figure 3.55 Experimental apparatus [58].

247

248

High Temperature Gas-cooled Reactors

Table 3.6 Main specifications of experimental apparatus [58]. Pressure: Temperature: Dimension and material: Reactor core flow passage: Graphite pipe Ceramic pipe Guide pipe High temperature plenum: Core barrel (inner barrel): Top cover: Pressure vessel: Outlet pipe:

Inlet pipe: Total height of test apparatus:

B0.27 MPa B1200 C

IG-110 (Toyo Tanso), 40 mm-φI.D. 3 70 mm-φO.D. 3 800 mm Al2O3 (Kyocera), 40 mm-φI.D. 3 70 mm-φO.D. 3 400 mm Al2O3 (Kyocera), 73 mm-φI.D. 3 85 mm-φO.D. 3 1250 mm Al2O3 (Kyocera), 400 mm-φI.D. 3 460 mm-φO.D. 3 150 mm Stainless steel, 800 mmφ 3 1519 mm Carbon steel, 460 mm in radius Carbon steel, 920 mm-φI.D. Al2O3 (Kyocera), 69.3 mm-φI.D. 3 80 mm-φO.D. 3 315 mm Stainless steel, 69.3 mm-φI.D. 3 76.3 mm-φO.D. 3 260 mm (Ultra-sonic anemometer) Stainless steel, 69.3 mm-φI.D. 3 76.3 mm-φO.D. 3 1100 mm 3750 mm

the reactor core simulator, and other structures were continuously measured at the inlet and outlet pipes, the high temperature plenum, the top space, and the annular passage between the inner barrel and the pressure vessel as shown in Fig. 3.55. The gas mixture in the apparatus was sampled at five positions as indicated in Fig. 3.55. The density of the gas mixture, the mole fraction of O2, CO2, and CO was measured using a gas analyzer (YOKOGAWA: density—Vibro gas analyzer DG-8; O2—Electrochemical analyzer 6234; CO2 and CO—Infrared rays analyzer IR21). Considering the errors induced by the thermocouples, the scanner junction, and the DVM accuracy, the entire accuracy of the temperature measurement was within 6 3 C. The relative uncertainties in the density of the gas mixture and in the mole fraction of O2, CO, and CO2 were found to be 6 4%, 6 2.3%, 6 4.4%, and 6 4.4%, respectively. The uncertainty in the inlet velocity was estimated to be 6 7%. The density and mole fraction changes in the apparatus were analyzed using FLOWGR. The apparatus is modeled as a network consisting of one-dimensional stream tubes. Fig. 3.56 shows the network model of the apparatus. At the beginning of the experiment, the apparatus was filled with helium gas. When the temperature of the gas and the reactor core simulator reached a steady state, the valves located at the ends of the inlet and outlet pipes were opened at the same time to simulate the pipe rupture accident condition. The parameters of the stream tubes are listed in Table 3.7. In this analysis, the molar-based formulation was used instead of the mass-based formulation. In the preliminary calculations, the two sets of formulations were checked by calculating

T 10

T7 7

5

Flow paths 1. Outlet pipe-1

12

18

16

T 15 T 16

T 17

T3 T2

T1

6. Peripheral graphite pipe 7. Peripheral ceramic pipe

T8

8. Pathway to insulator

3

9. Thermal insulator 10. Top cover-1 11. Top cover-2 12. Annular passage-1

2

13. Annular passage-2

1

15. Bottom cover-2 16. Coaxial pipe 17. Inlet pipe-1

P0T0

18. Inlet pipe-2

T: Temperature

Figure 3.56 Network model of the experimental apparatus [58].

4. Central graphite pipe

14. Bottom cover-1

T 18

P: Pressure (P′0>P0)

3. Plenum 5. Central ceramic pipe

150

406

14 T 14 15

4

T4 8

17

T9

T6

13

T 13

9

800

2. Outlet pipe-2

6

1431

T 12

500

T5

535

T 11

1471

10

11

Level

Pc0T0

Table 3.7 Parameters of stream tubes [58]. Path number

Path name

Length (m)

Equiv. diameter (m)

Wall temp. ( C)

C/O2 reaction

CO/O2 reaction

Vertical angle (degree)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18

Outlet pipe-1 Outlet pipe-2 Plenum Central graphite pipe Central ceramic pipe Peripheral graphite pipe Peripheral ceramic pipe Pathway to insulator Thermal insulator Top cover-1 Top cover-2 Annular passage-1 Annular passage-2 Bottom cover-1 Bottom cover-2 Coaxial pipe Inlet pipe-1 Inlet pipe-2

1.15 0.47 0.17 0.8 0.5 0.8 0.5 0.03 1.4 0.29 0.2 0.76 0.76 0.24 0.15 0.23 0.5 0.5

0.0693 0.0693 0.3767 0.04 0.0484 0.0693 0.0838 0.0417 0.7671 0.5919 0.8697 0.4543 0.4543 0.5155 0.3132 0.2086 0.0693 0.693

1428 28700 700900 908908 900200 908908 900200 900900 900900 104104 104225 225129 12933 2929 2914 1414 1414 1414

3 3 3 O 3 O 3 3 3 3 3 3 3 3 3 3 3 3

3 O O O 3 O 3 3 3 3 3 3 3 3 3 3 3 3

0 0 0 0 0 0 0 0 0 90 180 180 180 90 180 180 90 180

R&D on components

251

the process of air ingress in a reverse U-shaped tube with the graphite oxidation reaction. It has been found that fairly good agreements are obtained between the calculated results of the two sets of formulations with respect to the concentration changes of each gas species and the onset time of the natural circulation of air. It has also been found that the computations by the mass-based formulation take two to three times as much time as those by the molar-based formulation. In the case of a reverse U-shaped tube model, the matrix structure of the simultaneous linear systems is the simple tri-diagonal type. Thus the linear systems are efficiently solved by direct inversion. For the model, which has a complex network with many stream branches as shown in Fig. 3.56, the discretized equations are expressed by sparse matrices having a large bandwidth. As it is expected much more computation time will be required, the molecule-based formulation was adopted and the GaussSeidel iteration method was used for this analysis. Figs. 3.573.60 show an example of the measured density and mole fraction changes of O2, CO2, and CO at various measuring points. The average temperature of the reactor core simulator is about 900 C. The mole fractions of CO and CO2 in the plenum and the top cover gradually increase by molecular diffusion and the very weak natural circulation until the onset of the natural circulation of air. The first stage of the accident lasted for about 5 days in the experimental apparatus which simulates the HTTR. From this study, it became clear that air ingress process is completely different in the first and second stages. The important mechanisms of the air ingress in the first stage of the accident are molecular diffusion and the very weak natural circulation of the gas mixture. It also became clear that the duration of the first stage period is fairly long.

1.4 Average temperature (exp.) Outlet pipe (gas) =12°C High temp. plenum (gas) =856°C Graphite pipes (wall) =908°C Topspace (gas) =104°C Annular passage (gas) =102°C Bottom space (gas) =33°C

Density of the gas mixture (kg/m3)

1.2 1

Outlet pipe

0.8 Bottom cover 0.6

Annular passage

0.4 Top cover 0.2 Plenum 0 0

24 48 72 96 Elapsed time after the simulated pipe rupture (h)

Figure 3.57 Time-dependent behavior of the gas mixture density [58].

120

144

252

High Temperature Gas-cooled Reactors

(×10–2) 20

Average temperature (exp.) Outlet pipe (gas) = 12°C High temp. plenum (gas) =856°C Graphite pipes (wall) = 908°C Top space (gas) = 104°C Annular passage (gas) = 102°C Bottom space (gas) = 33°C

Mole fraction of O2

15

Outlet pipe

Plenum 10 Bottom cover Top cover annular passage

5

0 0

24

48 72 96 120 Elapsed time after the simulated pipe rupture (h)

144

Figure 3.58 Time-dependent behavior of the mole fraction of oxygen [58]. (×10–2) Average temperature (exp.) Outlet pipe (gas) = 12°C High temp. plenum (gas) =856°C Graphite pipes (wall) = 908°C Top space (gas) = 104°C Annular passage (gas) = 102°C Bottom space (gas) = 33°C

Mole fraction of CO2

10

8

6

Plenum

Top cover

4

Annular passage bottom cover

2 Outlet pipe 0

0

24

48 72 96 120 Elapsed time after the simulated pipe rupture (h)

144

Figure 3.59 Time-dependent behavior of the mole fraction of carbon dioxide [58].

The numerical results compared with the experimental ones are also shown in Figs. 3.573.60. As can been seen from these figures, the onset time of the natural circulation of air is fairly well reproduced and considerable agreements between the experiment and calculation are obtained for the density of the gas mixture and the mole fractions of each gas species.

R&D on components

253

(×10–2) Average temperature (exp.) Outlet pipe (gas) = 12°C High temp. plenum (gas) =856°C Graphite pipes (wall) = 908°C Top space (gas) = 104°C Annular passage (gas) = 102°C Bottom space (gas) = 33°C

Mole fraction of CO

8

6

4 Top cover 2

0

Annular passage bottom cover

Outlet pipe plenum

0

24

48 72 96 120 Elapsed time after the simulated pipe rupture (h)

144

Figure 3.60 Time-dependent behavior of the mole fraction of carbon monoxide [58].

References [1] S. Ueta, et al., Fuel performance under continuous high-temperature operation of the HTTR, J. Nucl. Sci. Technol. 51 (2014) 13451354. [2] S. Saito, et al., Design of High Temperature Engineering Test Reactor (HTTR), Japan Atomic Energy Research Institute, JAERI-1332, 1994. [3] K. Sawa, et al., Analytical method and result of off-site exposure during normal operation of High Temperature Engineering Test Reactor, Energy 16 (1991) 459470. [4] K. Sawa, et al., Safety criteria and quality control of the High Temperature Engineering Test Reactor fuel, Nucl. Eng. Des. 208 (2001) 305313. [5] K. Fukuda, et al., Research and development of HTTR coated particle fuel, J. Nucl. Sci. Technol. 28 (1991) 570581. [6] K. Fukuda, et al., Research and Development of HTGR Fuel, JAERI-M 89-007, 1989. [7] K. Minato, et al., Fission product palladium-silicon carbide interaction in HTGR fuel particles, J. Nucl. Mater. 172 (1990) 184196. [8] Y. Kurata, et al., The effect of heat treatment on density and structure of SiC, J. Nucl. Mater. 92 (1980) 351353. [9] G.T. Goodin, Accident condition performance of fuels for high-temperature gas-cooled reactors, J. Am. Ceram. Soc. 65 (1982) 238242. [10] H. Nabielek, et al., The performance of high-temperature reactor fuel particles at extreme temperatures, Nucl. Technol. 84 (1989) 6281. [11] T. Ogawa, et al., A model to predict the ultimate failure of coated fuel particles during core heatup events, Nucl. Technol. 96 (1991) 314322. [12] K. Sawa, et al., Fabrication of the first-loading fuel of the High Temperature Engineering Test Reactor, J. Nucl. Sci. Technol. 36 (1999) 683690. [13] S. Kato, et al., Fabrication of HTTR First Loading Fuel, IAEA-TECDOC-1210, 2001.

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[14] S. Ueta, et al., Development of high temperature gas-cooled reactor (HTGR) fuel in Japan, Prog. Nucl. Energy 53 (2011) 788793. [15] K. Minato, et al., Improvements in quality of as-manufactured fuels for hightemperature gas-cooled reactors, J. Nucl. Sci. Technol. 34 (1997) 325333. [16] T. Nishihara, et al., Excellent Feature of Japanese HTGR Technologies, JAEATechnology 2018-004, 2018. [17] K. Takamatsu, et al., High-temperature continuous operation of the HTTR, Trans. Atomic Energy Soc. Jpn. 10 (2011) 290300. [18] International Atomic Energy Agency, Fuel Performance and Fission Product Behavior in Gas-Cooled Reactors, IAEA-TECDOC-978, IAEA, Vienna, 1997, pp. 108-105. [19] S. Ueta, et al., Fuel and fission gas behavior during rise-to-power test of the High Temperature Engineering Test Reactor (HTTR), J. Nucl. Sci. Technol. 40 (2003) 679686. [20] Y. Inagaki, et al., R&D on thermal hydraulic of core and core-bottom structure, Nucl. Eng. Des. 233 (2004) 173183. [21] S. Maruyama, et al., Experimental studies on the thermal and hydraulic performance of the fuel stack of the VHTR Part I: HENDEL single-channel tests with uniform heat flux, Nucl. Eng. Des. 102 (1987) 19. [22] S. Maruyama, et al., Experimental studies on the thermal and hydraulic performance of the fuel stack of the VHTR Part II: HENDEL multi-channel test rig with twelve fuel rods, Nucl. Eng. Des. 102 (1987) 1120. [23] K. Kunitomi, et al., Thermal and hydraulic tests in HENDEL T2 supporting the development of the core bottom structure of the High Temperature Engineering Test Reactor, Nucl. Eng. Des. 108 (1988) 359368. [24] Y. Inagaki, et al., Hydraulic characteristics of coolant in the core bottom structure of the High-Temperature Engineering Test Reactor, Nucl. Technol. 99 (1992) 90103. [25] I. Ioka, et al., Thermal performance test of a coaxial double-tube hot-gas duct, Nucl. Technol. 105 (1994) 293299. [26] W.M. Kays, M.E. Crawford, Convective Heat and Mass Transfer, McGraw-Hill, New York, 1980. [27] M. Dalle Donne, E. Meerwald, Heat transfer and friction coefficient for turbulent flow of air in smooth annuli at high temperatures, Int. J. Heat Mass Transf. 16 (1973) 787809. [28] P. Bro¨ckerhoff, Behavior of thermal insulations for hot gas duct of high temperature gas cooled reactors, in: Proceedings of the ANS/ASME/NRC International Topical Meeting on Nuclear Thermal-Hydraulics, NUREG/CP-0014, 1980, pp. 23372351. [29] M. Hishida, et al., Thermal performance test of the hot gas duct of HENDEL, Nucl. Eng. Des. 83 (9) (1984) 1103. [30] S. Takada, et al., Research and development on passive cooling system, Nucl. Eng. Des. 233 (2004) 185195. [31] S. Takada, et al., Design and evaluation methods for a water cooling panel system for decay heat removal from a high temperature gas-cooled reactor, Heat Trans. Jpn. Res. 26 (1997) 159175. [32] A. Saikusa, et al., Advanced vessel cooling system concept for high-temperature gascooled reactors, Nucl. Technol. 118 (1997) 8996. [33] IAEA, Heat Transport and Afterheat Removal for Gas Cooled Reactors Under Accident Conditions, IAEA-TECHDOC-1163, 2001. [34] W.T. Sha, et al., A new approach for rod-bundle thermal-hydraulic analysis, Nucl. Technol. 46 (1979) 268280.

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[35] S. Howell, Thermal Radiation Heat Transfer, McGraw-Hill Kogakusha Ltd, 1972, p. 219. [36] A. Goldsmith, et al., Handbook of Thermo-physical Properties of Solids, Pergamon Press, 1961. [37] R.W. Thomas and de Vahl Davis, Natural convection in annular and rectangular cavities, Proc. Fourth Int. National Heat Tranf. Conf., Paris, Vol. 4 Paper NC 2.4, Elsevier Amsterdam. 1970. [38] M. Fishenden, O.A. Saunders, An Introduction to Heat Transfer, Clarendon Press, Oxford, 1950, p. 132. [39] Y. Shiina, et al., Natural convection heat transfer in hemisphere, J. Nucl. Sci. Technol. 25 (1988) 254262. [40] T.G. Theofanous, The in-vessel retention as a severe accident management strategy, in: Proceedings of the Eighth International Topical Meeting on Nuclear Reactor ThermalHydraulics, Kyoto, Japan, 1997, pp. 14. [41] O.A. Saunders, The effect of pressure upon free convection in air, Proc. R. Soc. Lond. A 157 (1936) 278. [42] Y. Inagaki, et al., R&D on high temperature components, Nucl. Eng. Des. 233 (2004) 211223. [43] I. Ioka, et al., Creep collapse of a heat transfer tube subjected to external pressure at high temperatures, Nucl. Eng. Des. 137 (1992) 259266. [44] Y. Kaji, et al., Estimation of creep buckling deformation under external pressure at elevated temperature, J. Pressure Vessel Technol. 118 (1996) 194197. [45] M. Kitagawa, et al., Lifetime test of a partial model of a high-temperature gas-cooled reactor helium-helium heat exchanger, Nucl. Technol. 66 (1984) 675684. [46] M. Futakawa, et al., Seismic test of a heat exchanger with a helically coiled tube bundle, Nucl. Technol. 118 (1997) 8388. [47] T. Iyoku, et al., Seismic response of the high-temperature engineering test reactor core bottom structure, Nucl. Technol. 99 (1992) 169176. [48] Y. Inagaki, et al., Thermal hydraulic study on a high-temperature gasgas heat exchanger with helically coiled tube bundles, Nucl. Eng. Des. 185 (1998) 141151. [49] Y.N. Chen, Flow-induced vibration and noise in tube-bank heat exchangers due to von Karman Streets, J. Eng. Ind. 90 (1968) 134146. [50] H.J. Connors, Fluidelastic vibration of tube arrays excited by cross flow, in: Proceedings of the Symposium on Flow Induced Vibration in Heat Exchangers, ASME Winter Annual Meeting, New York, December 1, 1970, pp. 4256. [51] M.J. Pettigrew, et al., Vibration analysis of heat exchanger and steam generators design, Nucl. Eng. Des. 48 (1978) 97115. [52] M.P. Paidoussis, Flow-Induced Vibration Design Guidelines, ASME, New York, 1981, p. 11. [53] A. Zukauskas, Heat transfer from tubes in crossflow, in: J.P. Harnet, T.F. Irvine (Eds.), Advances in Heat Transfer, Academic Press, New York, 1972, p. 142. [54] Y. Inagaki, Development of an in-service inspection technique for the intermediate heat exchanger tubes of the high-temperature engineering test reactor, Nucl. Technol. 104 (1993) 106117. [55] T. Takeda, M. Hishida, Study on diffusion and natural convection of two-component gases, Nucl. Eng. Des. 135 (1992) 341354. [56] T. Takeda, M. Hishida, Studies on molecular diffusion and natural convection in a multicomponent gas system, Int. J. Heat Mass Transf. 39 (1996) 527536. [57] M. Hishida, T. Takeda, S. Takenaka, Air ingress during the primary-pipe rupture accident of an HTGR, in: Proceedings of the 3rd JSME/ASME Joint International Conference on Nuclear Engineering, Tokyo, Japan, 1995, pp. 10931100.

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[58] T. Takeda, et al., Analysis of air ingress process during the primary-pipe rupture accident of the HTGR, in: Proceedings of the 3rd JAERI Symposium on HTGR Technologies, Japan Atomic Energy Research Institute, JAERI-Conf 96-010, 1996. [59] T. Takeda, Air ingress phenomena in a depressurization accident of the very-hightemperature reactor, Nucl. Eng. Des. 240 (2010) 24432450. [60] C.R. Wilke, A viscosity equation for gas mixtures, J. Chem. Phys. 18 (1950) 517. [61] R.K. Shah, A.L. London, Advances in Heat Transfer, Laminar Flow Forced Convection in Ducts, 143, Academic, New York, 1978, p. 78. [62] R.C. Reid, J.M. Prausnitz, T.K. Sherwood, The Properties of Gases and Liquids, third ed., McGraw-Hill, New York, 1977, pp. 3740. 226, 410414, 470474, 548565. [63] D.F. Fairbanks, C.R. Wilke, Diffusion coefficients in multicomponent gas mixture, Ind. Eng. Chem. 42 (1950) 471475. [64] R.E. Walker, N. deHaas, A.A. Westenberg, Measurements of multicomponent diffusion coefficients for the CO2-He-N2 system using the point source technique, J. Chem. Phys. 32 (1960) 13141316. [65] S.V. Patankar, Numerical Heat Transfer and Fluid Flow, McGraw-Hill, New York, 1980. [66] D.B. Spalding, A novel finite-difference formulation for differential expressions involving both first and second derivatives, Int. J. Numer. Methods Eng. 4 (1972) 551559.

Operation of HTTR

4

Yuji Fukaya, Minoru Goto, Hiroyuki Inoi, Etsuo Ishitsuka, Tatsuo Iyoku, Kazuhiko Kunitomi, Shigeaki Nakagawa, Tetsuo Nishihara, Hai Quan Ho, Akio Saikusa, Nariaki Sakaba, Hiroaki Sawahata, Taiju Shibata, Masayuki Shinozaki, Yukio Tachibana, Shoji Takada, Kuniyoshi Takamatsu and Daisuke Tochio Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan

The first criticality of the HTTR was achieved on November 10, 1998. After improvement of the reactor system and performance confirmation, the rise-to-power test was started in September 1999. The HTTR achieved the full power of 30 MW at a reactor outlet coolant temperature of about 850  C in December 2001, and the reactor outlet coolant temperature of 950  C was achieved in April 2004. After that, various operations such as high temperature operation at 950  C and safety demonstration tests were conducted until December 2010, and the HTTR has been suspended since the Great East Japan Earthquake in 2011. The problems during operation and the results of each operation are described as follows: 1. Unexpected incidents under construction and operation Two incidents during HTTR power-rise-tests, that is, temperature rise of the primary upper shielding and the core support plate, were described. Causes of the both incidents were gap flows of helium gas in structures. For the temperature rise of the primary upper shielding, countermeasures to reduce the gap flow, enhancement of heat release, and installation of thermal insulator were taken. For the temperature rise of the core support plate, temperature evaluations were carried out again considering the gap flow and design temperature of the core support plate was revised. By these countermeasures, the both temperatures were kept below their limits. 2. Characteristic test of initial HTTR core The results of core physics test in start-up and power-up of the HTTR were described. In the tests, the characteristics on the excess reactivity, the shutdown margin, the control rod worth, the reactivity coefficient, the neutron flux distribution, and the power distribution were investigated to verify the designed core performance and the required reactor safety. 3. Performance test of HTTR The results of the rise-to-power tests up to the full power were described. The control performance such the reactor power, the inlet and outlet coolant temperatures, the fission product behavior in the primary cooling system, the residual heat removal performance at reactor scram, etc. were investigated. 4. High temperature operation The long-term operations at 850 C operation for 30 days and 950 C operation for 50 days demonstrated that the HTTR is able to supply High Temperature Gas-cooled Reactors. DOI: https://doi.org/10.1016/B978-0-12-821031-4.00004-X © 2021 Elsevier Inc. All rights reserved.

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stable high temperature heat to a heat utilization system such as a hydrogen production system. The following two issues were validated using the HTTR burnup data: the effectiveness of rod-type burnable poisons at reactivity control and the whole core burnup calculation method. 5. 5. Safety demonstration test The tests on the inherent safety of HTGR by simulating various accidents such as the depressurization accident (loss of coolant accident), the reactivity insertion, and coolant flow reduction events, etc. were performed.

4.1

Unexpected incidents under construction and operation

4.1.1 Introduction In HTGRs, helium gas is used for the primary coolant. Helium gas leaks easily through narrow gaps by small pressure difference, transporting heat to unexpected parts. Thus it is an important issue for HTGRs to prevent local temperature rises by helium gas leakages. From the viewpoint of heat leakage, there were two incidents in the HTTR: One was a temperature rise of the primary upper shielding and the other was a temperature rise of the core support plate. This paper describes the works done to prevent temperature rises and countermeasures for the primary upper shielding and the core support plate.

4.1.2 Temperature rise of primary upper shielding 4.1.2.1 Outline of incident In 1997 nonnuclear heat up tests were carried out. When the primary coolant temperature became about 110 C by heat input from the gas circulators, the helium gas temperature around the control rod drive mechanisms inside the standpipes reached the alarm point of 60 C. At the same time, the temperature of the primary upper shielding reached about 75 C [1]. The structure of the standpipes and the related structures are shown in Fig. 4.1. Thirty-one standpipes on the top head dome of the reactor pressure vessel (RPV) penetrate the primary upper shielding. The primary upper shielding is a structure of concrete (grout), and carbon steel, that is the carbon steel box sustains weight as well as seismic load, and the top thick steel plate and the concrete function as shielding. Since the concrete of the primary upper shielding does not support load, the design temperature of the shielding concrete is set to 88 C. Heat from the RPV is conducted to the primary upper shielding through the standpipes. The standpipes and the top of the primary upper shielding are cooled by air flowing in the standpipe room as shown in Fig. 4.2. The air inlet temperature of standpipe room is about 15 C. The panel of the vessel cooling system (VCS) at the bottom of the shielding, which consists of many water-cooled tubes contained in a steel casing, also removes heat of the primary upper shielding. The water temperature can be adjusted between 17 C and 30 C all the year round.

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Figure 4.1 Structure of standpipes, RPV top head dome, primary upper shielding, etc. [1].

For the 16 standpipes for control rod drive mechanisms, a small amount of helium gas of about 25 C, called “purge gas,” from the helium purification system in the primary cooling system, comes in through the top of standpipe, purges control rod drive mechanisms, and cools the standpipes at the same time. The cause of the temperature rise of the primary upper shielding and the helium gas inside the standpipes was investigated by test and analysis. The primary coolant helium gas enters the bottom of the RPV and flows upward along its body. It turns around inside the top head dome, and then about 90% of the total primary coolant flows downward through 30 columns of fuel blocks.

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High Temperature Gas-cooled Reactors

Figure 4.2 Structure of inside of control rod standpipe before first countermeasure applied [1].

About 6% of the gas flows upward inside of the control rod guide sleeves and outside of the control rod guide tubes, changes its direction just under the heat insulators, and flows into the control rod guide tubes in order to cool the control rods as schematically shown in Fig. 4.2. The rest flows through gaps between the columns of blocks. In order to control the flow rate of the primary coolant in the control rod guide block column, there was one orifice of 11.3 mm in diameter, which produced a pressure drop between the plenum of the RPV top head dome, designated A in Fig. 4.2, and inside the guide pipe for the control rod support cable, designated B in the same figure. By the pressure difference, the primary coolant flows upward along the wall of the standpipe, changes its direction, and then flows downward through the graphite orifice and along the control rod support cable inside the guide pipe as shown in Fig. 4.2. The flow was named “bypass flow.” It was calculated that about 87% of the primary coolant for the control rod guide column goes through the orifice and about 13% bypasses the orifice. To reduce the temperature of the primary upper shielding, the two countermeasures were applied.

4.1.2.2 First countermeasures to reduce temperature 4.1.2.2.1 Countermeasure Countermeasure to prevent the bypass flow was considered and taken as shown in Fig. 4.3. The concept of the countermeasure is to minimize the pressure drop

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Figure 4.3 Structure of inside of control rod standpipe after first countermeasure applied [1].

between A and B in Fig. 4.2. For this purpose, square holes were bored on the side of the guide pipe for the control rod support cable and new holes for coolant paths were bored in the top plate of the control rod guide tubes as designated new openings. For controlling the flow rate of the primary coolant in the control rod guide block column, the multistage orifice was installed along the path for the control rod support cable instead of the orifice in Fig. 4.2. It was considered that the bypass flow could be prevented. At the same time, the gaps of the components for the control rod drive mechanism were sealed by synthetic rubber gaskets so that paths for leakage flow of helium gas are reduced and that most purge gas flows downward between the standpipe and the control rod guide sleeve as shown in Fig. 4.3. The purge gas works not only for cooling the control rod drive mechanism but also for preventing the bypass flow and cooling the standpipe.

4.1.2.2.2 Test result after first countermeasure The first countermeasure to the 16 control rod standpipes was finished in August 1997. Then nonnuclear heat up test, called the confirmation test 1, was conducted in September 1997 [2]. The results of the tests are shown in Fig. 4.4. The first countermeasure decreased the primary upper shielding temperature about 35 C when the reactor inlet coolant temperature was about 100 C. However, the shielding temperature rose to 64 C when the reactor inlet coolant temperature reached 164 C. Fig. 4.4 clearly shows that the shielding temperature exceeds the design temperature of 88 C at full power when the reactor inlet coolant temperature is 395 C.

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Figure 4.4 Test and analytical results of temperature of primary upper shielding [2].

Regarding the helium gas temperature inside the standpipe, its maximum temperature in the test was 32 C. The value is low enough compared with the alarm point of 60 C. The decrease in the helium gas temperature shows that the bypass flow is prevented. The test result indicates that the effect of the purge gas was much less than what we had expected, though the bypass flow is prevented. The reason why the purge gas is not effective is that the downward flow distribution is not uniform. The gap between the standpipe and the control rod guide sleeve, where the purge gas flows downward, is about 2 mm, the inside diameter of the standpipe is 485 mm, and the total length of the gap is 2600 mm. Differences in alignment of the standpipe and the control rod guide sleeve cause variation on the downward annulus flow. In the revised temperature analysis, the effective flow rate ratio of the purge gas that cools the standpipe was determined by trial and error so that nonnuclear heat up test results could be well reproduced by finite element method calculations. Then, the effective flow rate was used to predict the primary upper shielding temperature at full power.

4.1.2.3 Second countermeasures to reduce temperature 4.1.2.3.1 Countermeasure In the second countermeasure, heat removal by the VCS and heat release in the standpipe room were enhanced. The second countermeasure, installation of copper plates of about 3 mm thickness and heat insulators, is shown in Fig. 4.5.

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Figure 4.5 Second countermeasure to prevent temperature rise of primary upper shielding [1].

The upper part of the copper plate is composed of cylindrical plates around the standpipe, which are cooled by air of about 15 C in the standpipe room. The lower part of the copper plate contacts the vessel cooling panel at the bottom. Inside the panels are water-cooling tubes, which cool the bottom of the primary upper shielding directly and cool the standpipe side indirectly through the cylindrical copper plates. Between the standpipe and the lower copper plates are heat insulators made of microtherm. These prevent heat conduction and radiation from the standpipe. Heat insulators of kaowool were also put below the primary upper shielding around the standpipe to prevent radiation from the top head dome of the RPV to the standpipe.

4.1.2.3.2 Test result after installation of second countermeasure The nonnuclear heat up tests after installation of the second countermeasure, called the confirmation test 2 and the confirmation test 3, were conducted in 1998 and 1999, respectively [2]. Test results of both tests are also shown in Fig. 4.4. The temperatures in the confirmation test 3 were lower than those in the confirmation test 2. Concerning the two tests, the core of the HTTR is different. In the confirmation test 2, before fuel loading, the core consisted of dummy graphite blocks instead of fuel blocks. In the confirmation test 3, after the first criticality, the regular fuel blocks were installed. The core with dummy graphite blocks has larger core

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High Temperature Gas-cooled Reactors

pressure drop than the core with regular fuel blocks. Larger core pressure drop increases coolant flow rate in the control rod guide block columns. The core pressure drop in the confirmation test 2 and the confirmation test 3 was about 24 and 5.5 kPa, respectively. The calculations showed that increase in the core pressure drop decreases the ratio of the purge gas flowing between the standpipe and the control rod guide sleeve, and increases the ratio which goes through the graphite orifice. Decrease in flow rate of the purge gas increases temperature of the standpipes, which increase temperature of the primary upper shielding. In the confirmation tests 2 and 3, primary coolant was heated to 195 C and 213 C, respectively. In the confirmation test 2, the shielding temperature was 54 C when the reactor inlet helium temperature was 195 C. Comparing the result of the confirmation test 1 with the confirmation test 2, the temperature of the primary upper shielding decreased about 16 C when the reactor inlet coolant temperature was 170 C. The temperature fall is the effect of the second countermeasure. Difference in the temperatures of the primary upper shielding between the confirmation tests 2 and 3 was caused by the difference in the core as mentioned above. In the confirmation test 3, the effect of the mass flow rate of the purge gas on the temperature of the primary upper shielding was investigated. From the result of the test, the flow rate of the purge gas per each standpipe at normal operation was determined.

4.1.2.4 Prediction and test results at full power Temperature analysis of the confirmation tests 2 and 3 was conducted to estimate the temperature of the primary upper shielding at full power operation when the reactor inlet coolant temperature is 395 C. The results of the temperature analysis are also shown in Fig. 4.4. Analytical result of the shielding temperature at the rated power of 30 MW is 84 C for the confirmation test 2 C and 67 C for the confirmation test 3, which are lower than the design temperature of 88 C. The predicted temperature of the upper primary shielding with the determined flow rate of the purge gas mentioned above is 75 C as shown in Fig. 4.4, which is also below 88 C. Test results at power-rise tests up to 30MWare also shown in Fig. 4.4. The temperature was about 84 C at 30 MW, which corresponds to reactor inlet coolant temperature of about 395 C. The test results were higher than calculated results with the determined flow rate of purge gas. However, it was confirmed that the temperature of the primary upper shielding does not exceed the design temperature of 88 C at full power operation.

4.1.3 Temperature rise of core support plate 4.1.3.1 Outline of incident In the power-rise tests up to 20 MW, the core support plate showed unexpected temperature rise. It was predicted that its temperature should be higher than the design temperature of 470 C at 30 MW. The core support plate is installed at the bottom of the HTTR core. Since it was impossible to repair the reactor internal

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structure, temperature and integrity evaluation were carried out to confirm the integrity of the core support plate at the predicted high temperature. The reactor internals consist of graphite and metallic core support structures and shielding blocks as shown in Fig. 2.37. Core support plates are metallic component located at the bottom of the core support structures. The graphite core support structures, such as permanent reflector blocks and hot plenum blocks, are piled up on the core support plates, forming hot plenum and supporting the weight of core. The core support plates consist of seven hexagonal plates made of stainless steel and Cr Mo steel, being surrounded by another plate made of Cr Mo steel. The hexagonal plate in the center has a large hole in which the hot gas duct of a main cooling system penetrates. Another hexagonal plate also has a hole in which the hot gas duct of an auxiliary cooling system penetrates. Seal elements are put on the gaps between the core support plates to reduce helium leakage flow. The temperature of the center core support plate was the highest. The relation between the temperatures and the reactor outlet temperature is shown in Fig. 4.6. It was predicted that the temperature of the center core support plate would exceed the design temperature of 470 C at 30 MW rated operation. The other core support plates showed lower temperature. Their temperatures were expected about 400 C at 30 MW. The temperatures were also measured at the high temperature test operation up to 20 MW. The temperatures were lower than those in the rated operation as shown in Fig. 4.6. The difference between the rated operation and the high temperature test operation is the core flow rate by about 20%. It was considered that the temperature of the core support plate is affected by the core flow rate, which controls the amount of the helium leakage flow in the core support structures.

Figure 4.6 Core support plate temperature up to 20 MW in rated and high temperature test operation [2].

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High Temperature Gas-cooled Reactors

4.1.3.2 Reevaluation of core support plate temperature In the core support structures, there are many gaps between graphite blocks. To reduce the helium leakage flow in the gaps, called “gap flow,” seal elements are placed on the gaps. Almost all helium flows from the hot plenum into the hot gas duct. However, pressure drop between the hot plenum and inside of the hot gas duct causes a small amount of helium flowing into gaps between graphite blocks from the hot plenum. Schematic of gap flow is shown in Fig. 4.7. A few amount of helium flows downward and goes below the core support plate and flows to the center of core support plate. Then, helium flows up along the wall of the hot gas duct, flows into the mainstream of the hot gas duct of main cooling system. The helium from the hot plenum is about 850 C or 950 C. The helium from the hot plenum heats up the blocks and the temperature of core support plate rises. The core support structures of the HTTR are similar to those of the HENDL T2 test section, which was a test facility of HTGR core internals. In the tests of HENDL T2 test section, temperature distribution in the core support structures was measured. No significant temperature rise occurred in the structure [3]. It was that the effect of the gap flow is not significant. The coolant flow rate in the HENDEL T2 test section is about 1/3 1/4 of the HTTR core flow rate. Driving force of the gap flow is a pressure drop between the

Figure 4.7 Schematics of gap flow in core support structure [2].

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hot plenum and the hot gas duct of the main cooling system. The driving force of the gap flow in the HENDEL T2 was estimated about 1/9 1/16 of that in the HTTR. The gap flow in the HENDEL T2 test section was not large enough to raise the temperature of core support plates. In the HTTR, the gap flow became large because of large driving force of gap flow caused by the much larger core flow rate. Reevaluations of the core support plate temperature were carried out considering the effect of gap flow. The amount of gap flow was evaluated by a flow network code. The temperature of the core support structures was calculated by three-dimensional finite element method model. Fig. 4.8 shows the calculation model. Half of the core support structure is

Figure 4.8 Three-dimensional calculation model of core support structures by FEM calculation [2].

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High Temperature Gas-cooled Reactors

modeled. The evaluation of the gap flow rate made it clear that a small amount of helium flows in the core support structures. The total amount of the gap flow is about 0.1% of the core flow rate and increases with increase in reactor power. Temperature analyses of the core support plate were conducted considering the calculated gap flow. The comparison of the measured and calculated temperature distribution of the core support plate up to 20 MW is shown in Fig. 4.9. In the figure, calculation results with and without the gap flow are shown. The analytical results show good agreement with the measured results at the outer region. In the center region, analysis results with the gap flow show good agreement with the measured results. Analysis results without gap flow show flat temperature distribution and are much smaller than the measured results in the center region. With increase in thermal power, temperature rise in the center region became larger. By the temperature analysis, the temperature of the core support plate at 30 MW was predicted to be about 470 C at 30 MW. The design temperature of the core support plate was revised from 470 C to 530 C considering margins. Integrity of the core support plate under the revised design temperature was confirmed by stress analysis. Temperatures of the core support plate in the power-rise tests up to 30 MW in rated operation are shown in Fig. 4.10. At the 30 MW, the temperature of the core support plate was about 470  C, which agreed well with the analysis result.

Figure 4.9 Measured and calculated temperature profile of core support plate in rated operation [2].

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Figure 4.10 Measured core support plate temperature in rated operation up to 30 MW [2].

4.2

Characteristic test of initial core

4.2.1 Introduction Core physics test in the startup and the power-up operation was planned in order to ensure core performance and reactor safety of the HTGR. The results were obtained for a core burnup of up to 5 GWD/t. In the test, the critical approach, the excess reactivity, the shutdown margin, the control rod worth, the reactivity coefficient, the neutron flux distribution, and the power distribution were measured and compared with the calculated results [4]. The excess reactivity of the HTTR should be designed to exceed the minimum required for operations considering the reactivity loss caused by increase of core temperature and increased fission products buildup due to burn-up, etc. It should be also designed to keep the control rod insertion depth into the core low to maintain an acceptable power distribution. To attain the outlet coolant temperature of 950 C, it is important to keep the maximum fuel temperature as low as ever possible. This is because the fuel temperature is limited below 1495 C during normal operation and below 1600 C during anticipated operational occurrences (AOOs), in order to avoid significant fuel failure during operation. Power distribution in the core was optimized by changing uranium enrichment of each fuel assembly and by adapting burnable poison of boron carbide so that the maximum fuel temperature never exceed the limits [5]. For safety reasons, the reactor cores must have a negative reactivity feedback. It is very important to keep negative reactivity coefficient during all operational

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conditions throughout the whole burnup period. The control rods system of the HTTR consists of 16 pairs of control rods (32 rods, made of boron carbide) and drive mechanisms. Seven pairs of control rods are inserted into the fuel region and nine pairs into the reflector region. During normal operation, the positions of all control rods, except the three outermost control rods, are arranged at the same height not in order to disturb the axial power distribution.

4.2.2 General description 4.2.2.1 Test condition The fuel loading was started in September 1998 and the first criticality was attained in November 1998. The fully loaded core was completed in December 1998. During the fuel loading period up to the full core, various tests were conducted under cold clean condition as the startup core physics tests. During the startup core physics test, the reactor power was limited below 30 W. The helium pressure in the core was 1 bar. The tests for the critical approach, the excess reactivity, the shutdown margin, the neutron flux distribution, and the control rod worth were conducted. As the results of the tests, it was confirmed that nuclear characteristics required for full power operation such as the excess reactivity, the shutdown margin, and the control rod worth were adequate. The data for verification of the nuclear calculation code system were obtained [6,7]. The experimental results were summarized as benchmark problems of HTTR’s startup core physics in the IAEA coordinated research program [8]. The power-up tests ware started in April 2000. Reactor power was raised stepby-step in order to not to overlook significant defects in reactor safety. At each power level, various tests were conducted not only to confirm core characteristics but also to ensure the property of thermal hydraulic, radiation shielding, etc. The full power operation with an outlet coolant temperature of 850 C was attained in December 2001. The control rod positions for each power condition and the reactivity coefficients were measured from 0 up to 30 MW. The power distribution measurement of fuel assemblies was carried out during reactor shutdown after the power-up tests.

4.2.2.2 Nuclear calculations Nuclear calculations were performed using a diffusion code system and a Monte Carlo code. In the diffusion code system, lattice cell calculations to generate group constants were carried out by the DELIGHT and the TWOTRAN code, based on the collision probability method and the discrete ordinate transport theory, respectively. The DELIGHT code is a one-dimensional lattice burnup cell calculation code, which was developed for the nuclear design of the HTTR with neutron cross sections based on the ENDF-B/IV nuclear data library. Core calculations were performed by the CITATION code based on three-dimensional diffusion theory.

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The diffusion code system was applied to calculate the excess reactivity, the shutdown margin, the reactivity coefficient, and the neutron flux distribution. The continuous energy Monte Carlo code MVP [9] was used to calculate the critical approach, the control rod worth, the power distribution using cross sections based on the nuclear data library JENDL-3.2 or 3.3. The three-dimensional geometrical model for the HTTR core was described in detail with each fuel compact and burnable poison in the fuel assembly.

4.2.3 Critical approach The critical approach of the HTTR was carried out by the fuel addition method in which fuel assemblies replacing dummy graphite blocks previously loaded into the core. The fuel assembly was loaded from the core periphery to the core center. The annular core, such as a layered annular ring, was formed with 18 fuel columns. The order of the fuel loading is shown in Fig. 4.11. One of the purposes of this test is obtaining nuclear characteristic data of the annular core which are one of the promising core types for future HTGR because the core is effective to decrease the fuel temperatures during loss of coolant accident (LOCA) [10]. The experimental data of the core contribute to the verification 30 Column core 24 Column core

19 Column core (first critical)

18 Column core (subcritical)

Experiment Calculation

Figure 4.11 Change in keff value on critical approach at room temperature using fuel addition method [4].

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of calculation codes for future HTGRs. The fuel loading process consists of three phases. In the first phase, six fuel columns were loaded consecutively; in the second phase, three fuel columns were loaded until 15 fuel columns, and in the third phase fuel loading was carried out by one column at a time to identify the number of fuel columns for the first criticality. In every loading phase, all control rods were withdrawn from the core after each fuel loading and inverse multiplication factor (1/M value) were measured to predict the first criticality. With 19 fuel columns-loaded core, the first criticality was achieved. After the first criticality, the fuel loading was carried out additionally to 30 fuel columns for the fully loaded core. Prior to the fuel loading, the number of fuel columns for the first criticality was predicted as 16 6 1 by the MVP code with the use of JENDL-3.2, in which keff value was overestimated. After the test, the first critical number of fuel columns was reevaluated with reviewing the amount of impurity in graphite blocks and using the newest JENDL-3.3 library. The calculated critical number of fuel columns was 18 6 1 as shown in Fig. 4.11. Though the discrepancy of keff values from the measurement still remained, the tendency of calculated keff curve increasing with the fuel column number shows reasonably good agreement with the test results. Decrease of the discrepancy of keff is important to the development of future HTGRs using an annular core.

4.2.4 Excess reactivity and shutdown margin 4.2.4.1 Excess reactivity A fuel addition method with correcting the interacted reactivity effect among control rods was applied for the determination of the excess reactivity. The increment in reactivity by adding fuel was measured by the inverse kinetic method at a 21, 24, 27, and 30-columns loaded core. The interacted reactivity effect due to the control rods patterns in this measurement induced negative reactivity. Each reactivity increment was revised with the following relation. Excess reactivity is the sum of the revised reactivity increments. ρex 5 Ri 5

X

ΔρVi Δρai

Δρm 1 Ri

(4.1) (4.2)

V a where ρex , Δρm 1 , Δρi , Δρi , and I are excess reactivity, increment in reactivity measured with the inverse kinetic method, increment in reactivity calculated on a condition that all control rods are fully withdrawn, increment in reactivity calculated for the actual control rods pattern of the measurement, and number of fuel columns, respectively. Table 4.1 shows the measured and calculated values of excess reactivity in each fuel columns. The excess reactivity for the fully loaded core was 12.0%Δk/k. The calculated excess reactivity of the fully loaded core agrees well with the measured

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Table 4.1 Excess reactivity [4]. No. of fuel column

Experiment (%Δk/k)

Calculation (%Δk/k)

18 19 21 24 27 30

20.9 1.5 4.0 7.7 10.7 12.0

0.8 2.7 5.4 9.1 11.8 12.4

Table 4.2 Shutdown margin [4]. Scheme

Experiment (%Δk/k)

Calculation (%Δk/k)

IK method

Rod drop method

246.7

243.7

242.9

249.0

28.6 234.3 242.9

One-step scram All control rods Two-step scram 1st: control rods in reflector region 2nd: control rods in fuel region All control rods (total)

212.1 234.2 246.3

one, although core with less fuel columns did not agree so well. It was confirmed that the excess reactivity was below the limit of 16.5%Δk/k and slightly exceeded 11.1%Δk/k as required for the operation of 660 effective full power days.

4.2.4.2 Shutdown margin During full power operation of the HTTR, a two-step scram is employed in order to not exceed the temperature limit of control rods in the fuel region. At first, control rods at the reflector region are inserted; second, control rods at the fuel region are inserted after the control rod temperature has decreased sufficiently. The shutdown margin is defined as the negative reactivity of the core into which all control rods were inserted. To determinate the shutdown margin, the simulation of the two-steps scram was conducted during the startup core physics test under the cold clean condition. The scram reactivity was measured continuously by the inverse kinetic method. Table 4.2 shows the scram reactivity of reflector control rods and all control rods including the measurement results by the rod drop method. The calculated values by the diffusion code show good agreements with the measured ones in spite of the large negative value. It was confirmed that the shutdown margin exceeded the limit of 1%Δk/k even for the most severe cold clean condition and the stuckrod criterion.

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As the result, the measured shutdown margin was rather larger than the requirement because the shutdown margin of the HTTR is designed conservatively considering the calculation uncertainty and the cases where the control rods are removed from the core on a refueling and so on.

4.2.5 Control rod characteristics 4.2.5.1 Control rod worth on zero power condition Reactivity worth of each control rod was continuously measured by the inverse kinetic method from full-insertion to full-withdrawn level. The full-insertion and full-withdrawn levels correspond to 0 and 4060 mm, respectively, in axial distance from the bottom of fuel region as shown in Fig. 4.12. The control rod worth calculated by the MVP code agrees well with the measured values. Reactivity addition rate is one of the safety requirements of the HTTR and was calculated from a differential worth curve using the maximum withdrawing speed. It was confirmed that it is below the limit of 2.4 3 1024 Δk/k/s.

Experiment Calculation

Reflector or block

Fuel region

Reflector or block

Figure 4.12 Differential control rod worth for zero power operation [4].

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4.2.5.2 Control rod position versus reactor power The control rod position is defined as the distance from the fully inserted control rod. Fig. 4.13 shows the relation between the reactor power and the control rod position. The control rod position at full power operation was about 2850 mm, which corresponds to the control rod insertion depth of 50 mm into the fuel region. The control rod insertion depth at full power operation must be low so as not to disturb the axial power distribution significantly. It was confirmed that the control rod insertion depth was within the expected range of the design. The calculated values by the MVP code agree well with the measured ones within 5 mm except at 30 MW. The discrepancy at full power operation might be caused by the core temperature used for the calculation because the temperature, which is based on the conservative fuel temperature calculation, tends to be higher than the actual core temperature.

4.2.6 Reactivity coefficient 4.2.6.1 Temperature coefficient The reactivity coefficient of the HTTR is dominated by the temperature coefficient of the fuel and the reflector. The temperature coefficient was determined from the reactivity difference caused by a change of core temperature. The core temperature was increased at constant reactor power by reducing the heat removal in the secondary coolant system. The reactivity was estimated from the control rod worth curve after the core temperature was increased. Fig. 4.14 shows the results of the isothermal temperature coefficient measured up to 20 MW. Although the absolute values of the temperature coefficient decrease as the core average temperature increases, it was confirmed that all the measured values are negative. The measured values agree well with the

Figure 4.13 Relation between reactor power and the control rod position [4].

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Figure 4.14 Relation between temperature coefficients and average coolant temperature [4].

calculated ones for average coolant temperature beyond 400 C. Below 400 C, the experimental errors are very large and the coefficients do not agree so well.

4.2.6.2 Power coefficient The power coefficient consists of temperature coefficient and xenon reactivity worth. Reactivity change was estimated from the control rod position for each power level. Fig. 4.15 shows the results of the power coefficient measured from 0 to 30 MW. Although the absolute values decrease as the power increases, the power coefficient was confirmed to be negative for any power. Calculated values by the diffusion code agree well with the experiment for a power less than about 10 MW. The discrepancy between the measured and calculated values becomes larger beyond a core power of 10 MW.

4.2.7 Neutron flux and power distribution 4.2.7.1 Neutron flux distribution The axial neutron flux distribution for three irradiation columns in the reflector region is measured by fission chambers. The fission chambers were axially traversed by manually from the bottom of the fuel region to the upper surface of the second top reflector block. The output of fission chamber was affected by small reactivity change caused by the movement of the fission chamber. The count rate at each step was normalized with the count rate of the other fission chamber fixed at a certain position. Fig. 4.16 shows the measured and the calculated axial neutron flux distribution for a reactor power of 27 W. The control rod position at the measurement was about

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Figure 4.15 Relation between power coefficient and reactor power [4].

Experiment Calculation

Reflector or block

Fuel region

Reflector or block

Figure 4.16 Axial neutron flux distribution for zero power operation [4].

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1780 mm. The neutron flux distribution was calculated with the diffusion code. The measured distribution of the fission reaction rate obtained with the fission chamber is in good agreement with the calculation. From the results, it was confirmed that the calculation accuracy of neutron flux in the reflector region was satisfactory and the measurement provided fundamental data for irradiation tests.

4.2.7.2 Power distribution The power distribution of the core is an important property determining the maximum fuel temperature. Hence, the axial and radial shapes of the power distribution are designed to be exponential and flat, respectively, to keep the maximum fuel temperature as low as possible. Therefore the measurement was conducted to provide a basis for determination of the actual power distribution. The experiment was performed by measuring the gross gamma ray emitted from the fuel assembly when withdrawn from the core during reactor shutdown. The measured fuel assemblies were in service and their average burnup was about 4400 MWD/t. Prior to the measurement, a proportional relation between the gross gamma ray intensity and the power density of fuel assemblies was confirmed. A fuel assembly was withdrawn by the fuel handling machine up to the side where the Geiger Muller counter was mounted. The 20 fuel assemblies were measured and reloaded into the core after the experiment. The axial power distribution of the core is shown in Fig. 4.17. The axial distributions are normalized with the average of the core and the column, respectively. It was confirmed that the power profile of the HTTR agreed approximately with the expected one. Calculations to obtain the power distribution were performed by the MVP code. The calculated power distribution was integrated with the power at different power level from 9 to 30 MW considering the actual operation history. The calculation results show good agreement with the measurement. To ensure core performance and safety of the HTTR, the various tests were carried out at low burnup condition and nuclear characteristic data of the HTTR were obtained successfully. As for core performance, the obtained results from the core physics tests verified the expected characteristics for the critical approach, the neutron flux distribution, and the power distribution. The properties related to reactor safety were obtained from the measurement of the excess reactivity, the shutdown margin, the control rod worth, and the reactivity coefficient. All properties are satisfactory for a safe reactor operation with an outlet coolant temperature of 850 C.

4.3

Performance test

4.3.1 Introduction The construction of the HTTR started in March 1991 and the installation of major components was completed by March 1996. Comprehensive and functional tests for each system of the HTTR without fuel were started in October 1996, and several

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Experiment Calculation

Reflector or block

Fuel region

Reflector or block

Figure 4.17 Axial power distribution for 9 30 MW [4].

malfunctions requiring improvement were found. The improvements were finished by March 1998 before fuel loading. Fuel loading started in July 1998, and the HTTR attained the first criticality on November 10, 1998. The rise-to-power tests of the HTTR began in September 1999. The HTTR achieved the full power of 30 MW and the reactor outlet coolant temperature of about 850 C on December 7, 2001 in the rise-to-power tests [11]. Operation permit of the HTTR was obtained in March 6, 2002 from the Ministry of Education, Culture, Sports, Science and Technology, and after that, extensive tests such as safety demonstration test were started [12]. The maximum reactor outlet coolant temperature of 950 C was achieved in April 2004.

4.3.2 Major test items Major test items of the rise-to power tests are shown in Table 4.3. For safe and steady execution, the rise-to-power tests were conducted step by step, that is, they were divided into three phases of the power levels of 10, 20, and 30 MW. Test items of the rise-to-power tests can be categorized to tests for commissioning and for evaluating performance of the HTTR. The former test items are, for example, measurement of control rod reactivity worth, performance at abnormal transient (loss of off-site electric power test), radiation shielding performance, and measurement of radioactive material

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High Temperature Gas-cooled Reactors

Table 4.3 Major test items performed in rise-to-power tests [11]. Test item

10 MW

20 MW

30 MW

x x x

x x x x

x x x x

x

x

x

x x

x x

x x

x x x x x

x x x x

x x x x x

Commissioning test Control rod reactivity worth measurement at 30 kW Loss of off-site power Primary coolant and pressure measurement Radiation shielding performance Radioactive material concentration measurement in reactor building

x (30 kW)

Reactor and plant performance test Calibration of neutron instrumentation system to thermal reactor power Thermal-hydraulics in reactor core Heat exchanger performance in main cooling system Manual reactor scram Performance of vessel cooling system Fuel performance Measurement of impurity in primary coolant Thermal expansion performance of high temperature component

concentration in reactor building. The latter test items include performance of reactor control system, calibration of neutron instrumentation system to thermal power, performance of heat exchangers in the main cooling system, thermal expansion of high temperature components, thermal-hydraulics in reactor core, measurement of impurity in main cooling system, behavior of fuel and fission product. In the first phase test of up to 10 MW, core physics, radiation shielding performance, and performance of reactor control system were mainly investigated. Major objective of the second phase test of up to 20 MW was to investigate thermal properties such as performance of heat exchangers, thermal expansion of high temperature components, and thermal-hydraulics in the reactor core. The final third phase test of up to 30 MW was performed to confirm the overall function of the HTTR and to get the operation permit by undergoing the commissioning tests by the government. The rise-to-power tests at the rated operation mode were successfully completed in March 2002. The following subsection gives typical test results of the rise-to-power tests.

4.3.3 Heat balance of reactor cooling system Fig. 4.18 shows the measurement result of coolant temperature with design value in the main cooling system under the single-loaded operation. The reactor outlet

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Figure 4.18 Heat mass balance at single-loaded operation mode [11].

coolant temperature of 842 C is lower than the design value of 850 C because the heat loss of RPV turned out to be bigger than the design value by about 1% of rated reactor power of 30 MW. The heat loss of reactor cooling system is mainly due to the heat from the surface of RPV and the standpipe located at the top cover of the RPV top cover. The reactor outlet coolant temperature is adjustable to 850 C by fine adjustment of the primary coolant flow rate. The pressurized water temperature in the primary pressurized water cooler (PPWC) is lower than the design value because the overall heat transfer coefficient of PPWC is about 10% lower than the nominal design value. It was confirmed that the performance of PPWC was in the design range and the main cooling system was able to remove the heat of 30 MW generated in the reactor. The heat of about 97% generated in the reactor core is transferred through the PPWC to the pressurized water air cooler as a final heat sink. The reactor thermal output, which is used to calibrate the neutron instrumentation system, is evaluated from the heat removal of the pressurized water cooler, the VCS, and the auxiliary cooling system.

4.3.4 Heat exchanger performance In the single-loaded operation, only the PPWC is operated to remove the heat generated in the reactor core. The maximum thermal capacity of the PPWC is 30 MW in this operation mode. Table 4.4 shows the measurement data and evaluated

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High Temperature Gas-cooled Reactors

Table 4.4 Thermal performance of primary pressurized water cooler at 30 MW operation [11]. Operation mode

Amount of heat transfer (MW) Inlet helium gas temperature ( C) Outlet helium gas temperature ( C) Mass flow rate of helium gas (kg/s) Helium gas pressure (MPa) Inlet water temperature ( C) Outlet water temperature ( C) Mass flow rate of water (kg/s) Water pressure (MPa) KA value (kW/ C)

Rated and single-loaded operation Measurement

Design

28.8 834 390 12.6 4.0 106 146 625 3.4 63

29.9 843 384 12.6 3.9 135 175 625 3.5 71

thermal performances of the PPWC under the single-loaded operation with 30 MW. The thermal performance was evaluated by the following formula; KA 5

Q ΔTm

(4.3)

where K, A, Q, and ΔTm are the overall heat transfer coefficient, the heat transfer area, the amount of heat transfer, and the logarithmic mean temperature difference, respectively. The KA value evaluated by the measurement result was about 10% lower than that evaluated by design data. The evaluated result of thermal performance of the PPWC was in the range expected in the design and it was confirmed that the design method of heat exchanger in the HTTR is applicable. The difference of 10% between the design and the experiment is within the allowable limit estimated in the design of the PPWC, and it was confirmed that the main cooling system of the HTTR including PPWC can remove the heat of 30 MW generated in the reactor and achieve the high coolant temperature of 850 C and 950 C without overcooling.

4.3.5 Reactor control system performance The performance of the reactor control systems, that is, the pressurized water temperature control, the reactor inlet temperature control, and the reactor power control systems, was confirmed at a thermal power of 9 MW by giving some perturbations to the controllers. The control parameters such as a gain for proportional control action and a time constant of integral control action were determined in the performance test of reactor control system at this reactor power level. Fig. 4.19 shows the test result of the reactor inlet temperature control system comparing the calculation results. The condition of this performance test was a step-wise change of set value

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Figure 4.19 Performance of reactor inlet coolant temperature control system [11].

Figure 4.20 Performance of reactor power control system [11].

from 180 C to 170 C at a reactor power of 9 MW. After changing the set value of 10 C by step-wise in the reactor inlet temperature control system, the pressurized water temperature control system, which is the lower control system, increases the pressurized water flow rate in the air cooler to control the reactor inlet coolant temperature to control 170 C. According to the test result, the reactor inlet coolant temperature converged stably to the new set value of 170 C without any overshoot and oscillation. The calculation model can predict the performance of reactor inlet temperature control system because there was a good agreement between experimental and analytical results. After confirming all the control parameters of the reactor control system, the power-up test and the power-down test by the reactor control system were performed in the range from 30% to 35% of rated reactor power as shown in Fig. 4.20. From the experimental result, the reactor power increased gradually following the power-up rate of 1.5%/h given by the reactor power control system as a set value. The very slow power-up rate of 1.5%/h is for ensuring an integrity of the high temperature component such as an intermediate heat exchanger (IHX). By keeping the

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High Temperature Gas-cooled Reactors

Figure 4.21 Reactor operational trend up to full power [11].

power-up rate of 1.5%/h, the temperature change rate of the heat transfer tube made of Hastelloy XR in the IHX is limited under 15 C/h which is the requirement to ensure the integrity considering the creep and fatigue during the reactor power up. The reason why the reactor power increases with some time lag to the demand is that the reactor power control system has a dead band to the demand that is 0.4% (0.12 MW) of rated reactor power. Fig. 4.21 shows the measurement results of the reactor thermal power and reactor outlet coolant temperature in reaching to the rated power of 30 MW. The reactor outlet coolant temperature gradually reached to 850 C, because the reactor control system controlled the reactor power, reactor inlet coolant temperature, and coolant flow rate to constant values given from the heat balance of reactor cooling system.

4.3.6 Residual heat removal performance at manual reactor scram The manual reactor scram test from 30 MW was conducted to confirm the performances of the safety reactor shutdown and the residual heat removal by the auxiliary cooling system at the condition of an AOO postulated in the HTTR. In this test, the nine pairs of the control rods were inserted into the replaceable reflector region of the core just after the reactor scram. And then, the other seven pairs of the control rods were inserted into the active core region 40 min later prior to the decrease of the reactor outlet coolant temperature of 750 C after the reactor scram. The reactor power decreased and the reactor became subcritical after the reactor scram. Fig. 4.22 shows the transient behaviors of coolant flow rate, coolant temperature and component temperature. Two auxiliary helium circulators started up after the reactor scram, and the helium coolant flow rate of helium reached to the maximum value of about 1.8 kg/s within 20 s after the startup to the allowable limit of 1.2 kg/s. One of two auxiliary helium circulators coasted down 40 min later after the reactor scram and the coolant flow rate decreased to about 0.89 kg/s. The reason why such sequence is applied is to ensure the integrity of the control rods and the graphite components in the reactor core. The coolant temperature and component

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285

Figure 4.22 Temperature behavior during manual reactor scram test from 30 MW [11].

temperature of the plenum block and the flange of RPV top cover decreased very slowly after the reactor scram as shown in Fig. 4.22. Therefore it was confirmed that the temperature transient of the component is very slow as one of the inherent safety characteristics of HTGRs. And it was also confirmed that the auxiliary cooling system of the HTTR removed the residual heat from the reactor core continuously in the manual reactor scram test from up to 30 MW.

4.3.7 Thermal expansion performance of high temperature components The high temperature components such as the IHX, the PPWC, and the concentric hot gas duct are designed so as to compensate their thermal expansion by a floating support unit composed of constant hangers and oil snubbers as shown in Fig. 4.23. The floating support unit does not only compensate the thermal expansion but it also restrains the vibration under an earthquake and enables to minimize the size of the main cooling system with high temperature components. There are six measurement points with three directions measured by the strain gage-type displacement transducers in the main cooling system to confirm the thermal expansion performance [13]. The measurements of displacement of high temperature components were carried out at the reactor power levels of 9, 20, and 30 MW in the rise-topower tests. Fig. 4.23 shows the typical measurement result of displacement of high temperature component with a comparison of the analytical results by finite element method calculation codes (FEM codes). All the components in the main cooling system such as the PPWC, the IHX, the secondary pressurized water cooler, the helium gas circulator, and the concentric hot gas duct were modeled with beam elements in the FEM code to simulate the thermal expansion performance. The local rigidity at the connecting points between pipe and heat exchanger was considered in the analytical model. From the measurement result, the thermal expansion of all the components in the main cooling system was proportional to the thermal reactor power, that is primary coolant temperature, and there is no abnormal restriction to

286

High Temperature Gas-cooled Reactors

Figure 4.23 Experimental result of thermal expansion compared to analytical result [11].

Figure 4.24 Concentration of 88Kr in primary cooling system [11].

the component. The analytical results agreed well with the measurement results as shown in Fig. 4.23.

4.3.8 Fuel and fission product behavior Fuel and fission product behavior of the HTTR at several thermal power levels up to 30 MW was evaluated based on the measurement data by fuel failure detection system and primary coolant sampling system. Fig. 4.24 shows the primary coolant

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287

radioactivity of 88Kr, the primary coolant temperature, and the maximum fuel temperature to each reactor thermal power. The concentration of radioactivity was less than 5 3 1022 Bq/cm3, which is much smaller than the alarm level of 104 Bq/cm3 that is set as 1% of the total radioactivity in the reactor core. The release versus birth ratio of 7 3 1029 was much smaller than the design value of 5.4 3 1024. In the fabrication of the first-loading fuel of the HTTR, fabricated fuel compacts contained almost no wall-through failed particles and the average wall-through failure fraction was as low as 8 3 1025 [14].

4.4

High temperature operation

4.4.1 Introduction For commercialization of HTGRs, it is necessary to demonstrate long-term stable operation with a high outlet coolant temperature [15]. Thus a long-term high temperature operation test is planned. For safe and steady execution, the long-term high temperature operation was conducted step by step, that is, it was divided into two phases of outlet temperature levels of 850 C and 950 C. Continuous operation for 30 days in rated operation mode was performed in 2007. The fuel performance at retaining fission products (FPs) within the coating layer of the coated fuel particles (CFPs), reactor kinetics, and thermal expansion of the hot gas duct and heat exchanger was investigated. Continuous operation for 50 days in high temperature test operation mode was also performed in 2010. This added operation data at high temperature to establish the HTGR technical basis and to demonstrate the capability of high temperature process heat generation. In block-type HTGRs, reactivity is controlled with control rods (CRs) and burnable poisons (BPs). No other reactivity control systems are installed, such as chemical shims or recirculation flow controls in light water reactors (LWRs), which can be adjusted during operation. During full power operation, the CR insertion depth into the core should be kept shallow to keep the fuel temperature below the limit, because a large insertion depth leads to a significant disturbance of the power distribution in the core, and consequently the fuel temperature rises above the limit. As a consequence, there is relatively little control over the reactivity with the CRs, and reactivity control during the burnup period largely depends on the BPs. However, the effectiveness of the BPs at reactivity control has not been validated on HTGRs. It is necessary to validate the effectiveness of BPs at HTGR reactivity control. The HTTR adopts the rod-type BPs to control reactivity, and its burnup reached about 370 effective full power days (EFPD) during the long-term high temperature operation. Thereby, experimental data showing nuclear characteristics of burnup were obtained and used to validate the effectiveness of the BPs at reactivity control in the HTTR. Thus experimental validation of the effectiveness of BPs was performed using the HTTR burnup data. In addition, the whole burnup core calculation method, which is used to design the BPs, was also validated.

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High Temperature Gas-cooled Reactors

In this section, main test results of the long-term high temperature operation and the validation of the effectiveness of BPs and of the whole core burnup calculation method using the HTTR burnup data were described.

4.4.2 Main test results of long-term high temperature operation 4.4.2.1 30-Day continuous operation JAEA performed 30 days of continuous operation in rated and parallel-loaded operation mode in 2007 [16]. The behavior of the reactivity variation with burnup, fuel performance, thermal and nuclear power, and impurity concentration in the coolant was investigated. Fig. 4.25 shows the 30-day continuous operation history [17]. Reactor startup was performed on March 19th, 2007. The 30-day continuous operation at rated thermal power was held from March 27th to April 26th. Reactor shutdown was performed on May 3rd, 2007. Reactor power of the HTTR is usually controlled by the neutron flux signal from the power range monitoring systems. The gap between thermal power and the nuclear power changes with burnup. The reduction rate of thermal power was estimated to be about 12 kW/day. The thermal power is considered to decrease during rated operation as a result of consumption of heavy metals with burnup. The nuclear power shall be adjusted to the thermal power during longterm HTTR operation as needed. FPs are contained in the tristructural-isotropic (TRISO) CFPs but are partially released from failed fuel particles into the coolant as well as from uranium contamination in the fuel element matrix. As-fabricated and additional failed particles will be very few throughout the operation period. The concentration of FPs in the coolant indicates the core average coating failure fraction and the fuel matrix

Figure 4.25 30-Days continuous operation history [17].

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289

contamination fraction. Thus the failure fraction can be evaluated from the concentration of FPs. The FPs are measured by the helium sampling system and a Ge semiconductor analyzer. The release rate to birth rate (R/B) ratio of 88Kr was evaluated. The R/B of 88Kr is about 8.5 3 1029 at 30 MW of reactor power. Graphite and metallic core structures are corroded by impurities in the coolant such as water, CO, and CO2. These impurities are removed by the helium purification system. The concentration of impurities is automatically measured by the gas chromatography and moisture measuring system. The operational target values, measured concentrations of impurities, and the transient water concentration in the coolant are shown in Table 4.5 and Fig. 4.26. They are lower than the operational

Table 4.5 Operational target values of impurities in the coolant at rated operation [17]. Material

H2 H2O CO CO2 CH4 N2 O2

Target value at 850 C

Average

Max./min.

(vol. ppm)

(vol. ppm)

(vol. ppm)

7.5 0.5 7.5 1.5 1.25 0.5 0.1

0.06 0.15 0.09 0.01 0.01 0.02 0.03

0.19/0.01 0.18/0.11 0.16/0.05 0.02/0.01 0.01/0.01 0.01/0.01 0.01/0.01

Figure 4.26 Impurity concentration in coolant [17].

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High Temperature Gas-cooled Reactors

limits. Hastelloy XR is used as the material of the heat exchanger tube in the IHX. To maintain structural integrity and creep-deformation resistance, the coolant has controlled carburization and oxidation chemical conditions. The partial oxygen pressure is about 10 18, and the relative carbon activity is about 10 8, as shown in Fig. 4.27. The IHX is operated in suitable conditions for Hastelloy XR.

4.4.2.2 50-Day continuous operation Fig. 4.28 shows the 50-day continuous operation history. Reactor startup was performed on January 5th, 2010. The 50-day continuous operation at rated thermal power was held from January 22nd to March 13th. Power-down started in March 14th. Reactor shutdown was performed on March 21st, 2010. The most important characteristic of a HTGR fuel is its performance at trapping FPs in the enveloping layer. The primary coolant was measured during operation to observe the FP consistency, as shown in Fig. 3.8. The release of 88Kr is approximately constant, as it was during long-term rated operation. Compared to overseas data [18], the HTTR 88Kr release data are a few orders of magnitude lower. This result shows that the CFPs of the HTTR are excellent at confining FPs and have the highest performance in the world. Four kinds of impurity (H2, CH4, CO, and CO2) are produced from the graphite structure of the core internals with oxidization. It can be seen from the impurity data that oxidation of the graphite structure is within given design limits. Table 4.6 shows the temperature of the core internals during long-term high temperature operation compared to that during high temperature test operation in 2004. The location of each component is shown in Fig. 2.7. The temperature of the core restraint mechanism and the core support plate is below the given design limits. The temperature of the graphite structure is about the same as it was during test

Figure 4.27 Partial oxygen pressure and carbon activity [17].

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Figure 4.28 50-Days continuous operation history [17].

Table 4.6 Temperature of core internals [17]. Name of internals

Steel

Graphite

a

Core restraint mechanism Core support plate (central) Core support plate (periphery) Permanent reflector block (inner) Permanent reflector block (outer) Hot plenum block Lower block

Temperature reading on thermometer ( C)

Remarks

50-Day continuous operation

High temperature test operation in 2004

384 393

386 394

452

451

417 424

406 428

411 528

413 532

(9)b

419 468

422 471

(9)b

733 747 445 451

739 752 444 452

(3)b (3)b

Limit 450 Ca (11)b Limit 530 Ca (1)b Limit 505 Ca (3)b

Allowable working temperature for steel. Number of measuring points.

b

operation in 2004. Thus the measured temperature of the core internals is good agreement with the design value, which means that they will maintain their structural integrity.

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High Temperature Gas-cooled Reactors

The performance of the IHX was assessed in terms of a stable heat supply and the heat transfer performance. The secondary helium gas supply heat values during longterm high temperature operation were stable, as shown in Figs. 4.29 and 4.30. When heat transfer performance of the IHX decreases, the secondary helium gas temperature and reactor power decrease. The heat transfer performance of the IHX during longterm high temperature operation is compared with the results of high temperature test operation in 2004. It can be seen from the figure that the heat transfer performance of the IHX is the same as that during the test operation in 2004. Thus the IHX maintains excellent heat transfer performance from the beginning of operation.

4.4.3 Validation using high temperature engineering test reactor burnup data 4.4.3.1 Trend of change in control rod position Fig. 4.31 shows the changes in the CR positions with burnup for both zero power and full power in high temperature test operation mode. The experimental data were plotted for full power from 200 EFPD, at which the first full power hightemperature test operation was carried out [19], to the current 370 EFPD of burnup. For zero power, we plotted the experimental data from 0 EPFD to 300 EFPD. The experimental data for full power operation show that the CR position decreases with burnup, which suggests that the CR insertion depth increases with burnup. For zero power, meanwhile, the CR position is almost constant with burnup, which is different from the trend for full power. There is a different trend of the

Figure 4.29 Secondary helium gas supply heat values during long-term high temperature operation [17].

Heat exchanging performance (kW/K)

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293

120 100 80 60 High temperature test operation (2004) 50-day continuous operation (2010)

40 20 20

40

60

80

100

120

140

Log-mean temperature difference ( C) Figure 4.30 Heat transfer performance of IHX during long-term high temperature operation [17].

Figure 4.31 Changes in CR position with burnup [17].

changes in the CR positions with burnup between zero power and full power. The reactivity balances for zero and full power are expressed by Eqs. (4.4) and (4.5), respectively, through the burnup: ρFuel;Zero 5 ρBP;Zero 1 ρCR;Zero

(4.4)

ρFuel;Zero 5 ρBP;Zero 1 ρTemp 1 ρCR;Full

(4.5)

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High Temperature Gas-cooled Reactors

where ρFuel,Zero, ρBP,Zero, ρCR,Zero, ρCR,Full, and ρTemp are the reactivity of fuel for zero power, the reactivity compensated by BP for zero power, the reactivity compensated by CRs for zero power, and the reactivity compensated by CRs for full power and the reactivity loss due to the temperature increase from zero to full power conditions, namely, temperature reactivity defect. The curve for zero power in Fig. 4.31 shows that the CR insertion depth was almost constant from BOL to 300 EFPD. Hence, let the change of ρCR,Zero be ΔρCR,Zero when burnup increases infinitesimally, ΔρCR;Zero 5 0

(4.6)

where we assumed that the variations in the CRs’ worth with burnup that affect the changes in the CR position are negligibly small. No variations in the CRs’ worth with burnup have been observed in the HTTR data yet. Eq. (4.4) becomes the following equation using Eq. (4.6): ΔρFuel;Zero 2 ΔρBP;Zero 5 0

(4.7)

where ΔρFuel;Zero indicates changes of ρFuel;Zero when burnup increases infinitesimally. Meanwhile, the curve for full power in Fig. 4.31 shows that the CR insertion depth increased with burnup. Thus let a change of ρCR;Full be ΔρCR;Full when burnup increases infinitesimally, ΔρCR;Full . 0

(4.8)

where we used the same assumption for variation in the CRs’ integrated reactivity value as we did for zero power. Eq. (4.5) becomes the following equation using Eq. (4.8): ΔρFuel;Zero 2 ΔρBP;Zero 2 ΔρTemp . 0

(4.9)

where ΔρBP;Zero and ΔρTemp indicate changes of ρBP;Zero and ρTemp , respectively, when burnup increases infinitesimally. From Eqs. (4.7) and (4.9), the following is obtained: ΔρTemp , 0

(4.10)

Eq. (4.10) indicates that the reactivity loss due to raising the core temperature has decreased with burnup, which causes the different trend of the changes in the CR positions with burnup between zero and full power. In order to understand the reason for the decrease in the temperature reactivity defect, cell burnup calculations, in which the depletion equation is solved, were performed. Then, the temperature reactivity defect was calculated for each burnup

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295

step. The cell burnup calculations were performed using the averaged enrichment and temperature conditions of the HTTR. Subsequently, branch-off calculations were performed, in which the depletion equation was not solved and cell calculations were performed for each burnup step by using the composition obtained by usual cell burnup calculations. In the branch-off calculations, the number densities of 239Pu and 241Pu were defined to be zero for each burnup step, and the temperature reactivity defect was also calculated. Fig. 4.32 shows that the temperature reactivity defect obtained by the cell burnup calculations decreases about 2.5%Δk/k during the burnup period, while that obtained by branch-off calculations is almost constant. Thus the decrease in the temperature reactivity defect indicated by Eq. (4.10) is mainly caused by accumulation of plutonium. The effect of plutonium accumulation on the nuclear characteristics of HTGRs is larger than that on other thermal reactors such as LWRs. The reason for this is that the neutron spectrum in HTGRs is harder than that in other reactors because of the high temperature of HTGRs’ cores and the low-energy fission cross-section resonance of 239Pu and 241Pu. It is predicted that in the HTTR, the temperature reactivity defect decreases about 2.5%Δk/k during the burnup period. Thus treating the low-energy resonances of 239Pu and 241Pu properly is particularly important in the nuclear design of HTGRs.

4.4.3.2 Effectiveness of rod-type burnable poisons 4.4.3.2.1 Design philosophy of burnable poisons In the HTTR, fuel cannot be added into the core during operation, and accordingly enough fuel should be loaded into the core at the beginning of life (BOL). If the core contains only fuel, excess reactivity becomes very high, a situation that requires a large insertion depth of the CRs. The large insertion depth causes a

Figure 4.32 Effect of accumulation of plutonium on temperature reactivity defect [17].

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High Temperature Gas-cooled Reactors

significant disturbance to the optimized power distribution in the axial direction, and consequently the fuel temperature rises above the limit. Thus reactivity control is very difficult with only the CRs. In order to resolve this problem, BPs are loaded into the core in the HTTR. BPs are designed to flatten the excess reactivity changes with burnup in full power operation as shown in Fig. 4.33 [20] and maintain the optimized power distribution during the burnup period. BPs are specifically designed to allow a shallow CR insertion depth, above the middle of the first layer of the core, during the burnup period.

4.4.3.2.2 Validity of effectiveness of rod-type burnable poisons In order to validate the effectiveness of the rod-type BPs at reactivity control in the HTTR during high temperature test operation, it was investigated whether the BPs had functioned as designed, specifically focusing on whether the CR insertion depth into the core was kept above the middle of the first layer of the core during the burnup period. In the HTTR, each CR insertion depth is obtained from the CR position value. The CR position is defined as the distance from the bottom of the fuel region to the tip of the CR. In addition, the positions of the C-CR, R1-CRs and R2CRs used for control of reactivity during normal operation are the same within a range of 10 mm. It was investigated whether the averaged CR position remained at the designed position. Fig. 4.31 shows that in full power operation, the CR position has been kept above the middle of the first layer of the core from BOL to the current burnup of 370 EFPD. The extrapolated line using the experimental data obtained around 200 EFPD suggests that the CR insertion depth at the end of life is beyond the middle of the first layer of the core. However, the rate of increase of the CR insertion depth decreases with burnup, because of the plutonium accumulation. Thus the absolute

Figure 4.33 Change in excess reactivity of HTTR in full power operation when CRs are fully withdrawn [20].

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297

value of the slope of the CR position curve becomes smaller with burnup, and the CR position remains above the middle of the first layer of the core after 370 EFPD. Hence, for the HTTR full power operation, the CR position was confirmed to be kept above the middle of the first layer of the core during the burnup period, as expected. Thus the effectiveness of rod-type BPs at reactivity control in the HTTR was validated.

4.4.3.3 Whole core burnup calculations 4.4.3.3.1 Calculation method Whole core burnup calculations for the HTTR were performed in two steps. First, cell burnup calculations were performed with the SRAC/PIJ code [21] using the collision probability method to generate few group cross sections with burnup dependence. Second, the whole core calculations were performed with the SRAC/ COREBN code [21] based on diffusion theory for each burnup step using the few group cross sections. 1. Cell calculations In the cell burnup calculations to generate few group cross sections for the fuel blocks, the fuel block geometry was used as a unit cell instead of a fuel pin model. Fuel and BPs were treated as burnup materials, and the double heterogeneity effect caused by the CFPs was taken into consideration. Two homogenized macro cross sections with burnup dependence, one of which contained BPs and the other of which did not as shown in Fig. 4.34, were generated for each fuel block. In cell calculations for the other blocks, each block geometry was also used as the unit cell. In the cell calculations for the CR guide blocks,

Figure 4.34 Geometry model of whole core burnup calculation for HTTR [17].

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High Temperature Gas-cooled Reactors

the neutron absorber contained in the CRs was treated as a no-burnup material. Two homogenized macro cross sections, one of which contained the CRs and the other of which did not (Fig. 4.33), were generated. In the cell calculation for reflector and irradiation blocks, one homogenized macro cross section was generated. 2. Core diffusion calculations

A three-dimensional triangular mesh was used in the core burnup calculations. Each block composing the core was divided into 24 triangular meshes horizontally and into 14 meshes vertically.

4.4.3.3.2 Validity of whole core burnup calculations The validation of the whole core burnup calculations was performed by comparing the CR position, which changes with burnup, obtained from the calculations with that from the HTTR data. For the HTTR full power operation, the experimental data and the analytical result showed the same trend for the change in the CR position with burnup. They agree within 10 cm throughout the burnup period, which corresponds to about 1%Δk=k in reactivity. Meanwhile, for zero power, both the experimental data and the calculation results show that the CR position is almost constant throughout the burnup period. They agree within 15 cm, which corresponds to about 1.5%Δk/k in reactivity, during the burnup period. Hence, although there is still room for improvement, the validity of the whole core burnup calculation for the HTTR with the SRAC/COREBN code was confirmed by the experimental data.

4.5

Safety demonstration test

4.5.1 Introduction In the safety evaluation of the HTTR, its inherent safety as the HTGR and specific design features of the HTTR were taken into account. The categorization of the events to be evaluated was based on the “Guidelines for Safety Evaluation of LWR Power Plants” [22]. In other words, so-called design basis events were selected for AOOs as shown in accidents, major accidents, and hypothetical accidents. These events include the events selected for confirmation of the adequacy of the modification of scram set values for safety demonstration tests and for confirmation of integrity of test facilities such as irradiation capsules and test samples. The AOOs include conditions beyond the normal reactor operation resulting from a single failure or malfunction, or a single operator error anticipated to occur during the lifetime of the reactor facility, as well as one that may occur with a similar frequency as the above and result in unplanned operating conditions. The accidents are beyond the scope of the AOOs and should be determined for safety evaluation of a reactor facility from the viewpoint of the possibility of radioactivity release though the frequency is smaller. Major and hypothetical accidents were evaluated in order to judge the appropriateness of the reactor siting conditions based on the “Guidelines for Reactor Siting Evaluation” [23].

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Acceptance criteria for the HTTR that reflect the safety requirements for LWR power plants and take into account major features of HTGRs and the HTTR have been established. For example, the maximum fuel temperature is restricted to 1600 C, which was determined based on the test results, to avoid fuel failure during AOOs. Abnormal events to be postulated as AOOs and accidents were selected considering their frequencies of occurrence and based on the investigation of main causes that affect each item of the acceptance criteria identified for the HTTR, that is, (1) fuel temperature, (2) core damage, (3) temperature of the reactor coolant pressure boundary, (4) pressure at the reactor coolant pressure boundary, (5) pressure at the containment vessel boundary, and (6) risk of radiation exposure for the public. The initiating abnormal events were classified into similar event groups according to the “Guidelines for Safety Evaluation of LWR Power Plants” [22]. Then, the most severe events with respect to the acceptance criteria within each similar event group were selected as the representative postulated events. The main causes that affect the fuel temperature are increase of power and decrease of heat removal in the core. The increase of power is caused by reactivity addition, in which four event groups are postulated, namely (1) malfunction of the reactivity control system, (2) malfunction of the experimental facility, (3) increase of the primary coolant flow, and (4) increase of heat removal by the secondary cooling system. Two initiating events are considered to cause the malfunction of the reactivity control system, namely (1) abnormal control rod withdrawal and (2) abnormal control rod insertion. Abnormal control rod withdrawal is the more severe event with respect to the fuel temperature and is selected as the representative postulated event. The representative postulated events concerning other acceptance criteria are selected in the same way. The postulated events considered in the safety evaluation of the HTTR as AOOs and accidents are listed in Table 4.7 [12]. The safety demonstration tests are conducted to demonstrate inherent safety features of the HTGRs as well as to obtain the core and plant transient data for validation of safety analysis codes and for establishment of safety design and evaluation technologies of the HTGRs [12,24,25].

Table 4.7 Postulated events classified as AOOs [12]. AOO-1 AOO-2 AOO-3 AOO-4 AOO-5 AOO-6 AOO-7 AOO-8 AOO-9

Abnormal control rod withdraw under subcritical condition Abnormal control rod withdrawal during rated operation Decrease in primary coolant flow rate Increase in primary coolant flow rate Decrease in heat removal by secondary cooling system Increase in heat removal by secondary cooling system Loss of off-site electric power Abnormality of irradiation specimens and experimental equipment Abnormality during safety demonstration tests

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High Temperature Gas-cooled Reactors

4.5.2 High temperature engineering test reactor control system 4.5.2.1 Reactivity and reactor power control systems The HTTR core consists of thirty columns of fuel elements, seven columns of control rod guide blocks, replaceable reflector, and large-scale permanent reflector blocks. Sixteen pairs of control rods in the fuel and replaceable reflector regions of the core control the reactivity in the HTTR. A control rod drive mechanism drives each pair of control rods using an AC motor. At a reactor scram, electromagnetic clutches of the control rod drive mechanisms are separated and the control rods fall into holes of the control rod guide blocks by gravity at a constant speed, shutting down the reactor safely. In an unlikely event that the control rods insertion fails, reserved shutdown pellets made of B4C/C are dropped into the holes instead of the control rods. The reactor power control system and the reactor outlet coolant temperature control system are cascaded: the latter control system is superior to the reactor power control system as shown in Fig. 4.35. The signals from each channel of the power range monitoring system are transferred to three controllers using the microprocessors. In case there is a deviation between the process and set values, a pair of

Figure 4.35 Reactor power control device of the HTTR [12].

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control rods is inserted or withdrawn at the control rod at a speed from 1 to 10 mm/s according to the deviation. The relative positions of 13 pairs of control rods, except for 3 pairs of control rods used only for the scram, are controlled within 20 mm one another by the control rod pattern interlock to prevent any abnormal power distribution in the reactor. In case of a reactor scram, except for a depressurization accident, the control rods are inserted into the replaceable reflector region first, and then, the remaining control rods are inserted into the fuel region under the condition that the reactor outlet coolant temperature has descended lower than 750 C or 40 min has elapsed after a reactor scram. In the case of depressurization accident, all control rods are inserted into the core simultaneously.

4.5.2.2 Cooling system and plant control device Fig. 4.36 shows the cooling system and plant control device of the HTTR. The main cooling system operates at normal operation, and an auxiliary cooling system as well as a VCS, which operates to remove residual heat of the core after a reactor scram. The auxiliary cooling system and the VCS are engineered safety features. In commercialization of economically competitive HTGRs, it is necessary to eliminate the engineered safety features by establishing new safety philosophy based on the results of the HTTR safety demonstration tests. The elimination of the engineered safety features is helpful to reduce the construction cost of the HTGR. The main cooling system removes heat generated in the core and dissipates it to the

Figure 4.36 Cooling system and plant control device of the HTTR [12].

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High Temperature Gas-cooled Reactors

atmosphere by a pressurized water air cooler in the pressurized water-cooling system. The HTTR has two operation modes regarding use of heat exchangers for both of “the rated operation mode” and “the high temperature test operation mode.” At “the single-loaded operation mode,” only the PPWC is operated in the primary cooling system, whereas at “the parallel loaded operation mode,” both the intermediated heat exchanger and PPWC are operated, and the intermediated heat exchanger and the PPWC remove heat of 10 and 20 MW, respectively.

4.5.3 Safety demonstration test plan Table 4.8 shows the planned safety demonstration tests which are separated into the following six tests: 1. 2. 3. 4. 5. 6.

reactivity insertion test—control rod withdraw test, coolant flow reduction test—partial loss of coolant flow test, coolant flow reduction test—gas circulator trip test, loss of forced cooling test—all gas circulator trip test, all blackout test— VCS stop test, and depressurization test—simulation of LOCA.

Table 4.8 Planned safety demonstration tests and conducted tests. Planned test item

(1) Reactivity insertion test— control rod withdraw test

(2) Coolant flow reduction test— partial loss of coolant flow test (3) Coolant flow reduction test— gas circulator trip test

(4) Loss of forced cooling test—all gas circulator trip test (5) All black out test—vessel cooling system stop test (6) Depressurization test— simulation of loss of coolant accident (LOCA)

Conducted tests Test no.

Year

Reactor power (MW)

Flow rate (%)

CR-1 CR-2 CR-3 CR-4 GC-5 GC-7 GC-9 GC-1 GC-2 GC-3 GC-4 GC-6 GC-8 GC-10

2002 2003 2004 2006 2004 2006 2007 2003 2003 2004 2004 2006 2007 2010

9 15 18 24 18 24 30 9 9 18 18 24 24 9

100 (GCs) 100 (GCs) 100 (GCs) 100 (GCs) 98 (GCs) 98 (GCs) 98 (GCs) 67 (GCs) 33 (GCs) 67 (GCs) 33 (GCs) 33 (GCs) 33 (GCs) 0 (GCs)

GC/VCS-1

2010

9

0 (GCs), 50 (VCS)

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In the reactivity insertion test, a central pair of control rods is withdrawn and a reactivity insertion event without scram will be simulated. The coolant flow reduction test is composed of the partial loss of coolant flow test and the gas circulators trip test. In the partial loss of coolant flow test, the primary coolant flow rate is slightly reduced by the control system of the primary coolant flow rate with the reactor outlet coolant temperature control system being operated. In the gas circulators (GCs) trip test, the primary coolant flow rate is reduced to 67% and 33% of rated flow rate by running down one and two out of three gas circulators at the PPWC without a reactor scram, respectively. In the loss of forced cooling test, all the three gas circulators at the PPWC are run down without a reactor scram, and loss of forced cooling will be simulated. In the all blackout test, the VCS, which is an engineered safety feature and has two independent systems, is shut down in addition to the run-down of all the three gas circulators at the PPWC. This test simulates an accident in which all the cooling system is run down without a reactor scram. In the depressurization test, the primary coolant pressure is reduced by removing primary coolant of helium gas to storage tanks in addition to the stop of all the gas circulators, simulating a LOCA. The conducted tests are also shown in Table 4.8. The tests on (4) (6) have been suspended after the Great East Japan Earthquake in 2011.

4.5.4 Analysis code and model The steady-state and transient behaviors of the reactor and plant of the HTGR can be evaluated using the ACCORD code [26]. The major characteristics of this code are as follows: 1. The plant system can be analyzed for long term after an event starts by modeling the heat capacity of the reactor core. 2. Thermal hydraulics for each component such as the reactor, heat exchangers, and concentric hot gas ducts can be analyzed by separating the heat transfer calculation for a component from the fluid flow calculation for helium and water.

The ACCORD code consists of modules for nuclear calculation, heat transfer calculation in the reactor, heat exchangers and piping, fluid flow calculation of helium and water, and control system and safety protection system of the HTTR. Fig. 4.37 shows a calculation scheme of the ACCORD code. Nuclear characteristics are evaluated by conventional point kinetics model with six delayed neutron groups. The thermal power is calculated by a balance of feedback reactivity due to fuel and moderator temperatures in the reactor core taking into account additional reactivity caused by inserting and withdrawing the control rods. Decay heat after the reactor scram is estimated according to Shure’s formula and decay heat of actinide. The reactor core is simulated by one channel model for one fuel rod considering thermal conduction through components and heat transfer by helium. Each heat exchange is simulated by one channel model with one heat transfer tube. Helium and water flow are approximated by one-dimensional flow network model including flow line and pressure point for calculating the flow rate and pressure. Improvement of the

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Figure 4.37 Calculation scheme of ACCORD code [26].

reactor core model to multichannel was carried out utilizing results of the safety demonstration tests. The control system module simulates proportional and integral control applied to the HTTR control system for reactor power, inlet and outlet temperatures of the reactor, primary and secondary helium flow rate, etc. The safety protection system module simulates the HTTR scram signal system. It observes the lags of time until the occurrence of the scram signal and until the startup of the auxiliary cooling system.

4.5.5 Reactivity insertion test 4.5.5.1 Objective and test procedure The reactivity insertion test demonstrates that rapid increase of the reactor power following withdrawal of the control rods is restrained only by the negative feedback of reactivity without operating the reactor power control system, and the temperature transition of core component is slow over the event due to its large heat capacity. The test data obtained are used for the development and validation of safety evaluation codes for HTGRs. In the reactivity insertion test, the central pair of control rods out of 16 pairs in the core is withdrawn with the reactor power control system being blocked. The central pair with the maximum control rod worth was chosen for the test. In this test, power supply for driving all the control rods, except the central pair, is cut off to prevent withdrawal of control rods by mistake. In addition, the normal value of

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20 mm for the control rod pattern interlock is modified to 50 mm, which corresponds to the limitation of continuous withdrawal of control rods for the HTTR. The cut-off of power supply driving the control rods and modification of the control rod pattern interlock is done automatically upon selection of the test mode switch. The reactor power and temperature transients are measured by ionization chambers and thermocouples, respectively. Major measurement points and instrumentation are shown in Fig. 4.38.

4.5.5.2 Test results The results of control rod withdrawal test at the reactor power of 15 MW are compared with analysis results as shown in Figs. 4.39 4.41 [24]. Fig. 4.39 shows the transient of the reactor power and control rods position. Positive reactivity is inserted when the central pair of control rods is withdrawn. Although the reactor power increases because the positive reactivity is not compensated by other control rods insertion, the reactor power decreases due to negative feedback effect with increase of core temperature. As shown in Fig. 4.39, the reactor power rises up to

Figure 4.38 Measurement points and instrumentation [12].

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High Temperature Gas-cooled Reactors

Figure 4.39 Transient of reactor power and the control rod position during control rods withdrawal test [24].

Figure 4.40 Transient of the fuel and moderator temperatures during control rods withdrawal test [24].

about 16.7 MW from the initial reactor power of 15.2 MW and decreases afterward. Finally, the reactor power approaches to about 15.5 MW as the positive reactivity by control rod withdrawal compensates with the negative one by core temperature increase. The average fuel temperature rises up to about 505 C from the initial one of 498 C as shown in Fig. 4.40. The fuel temperature gets to maximum about 30 s after the reactor power reaches maximum, due to heat capacity of fuel compacts and time lag of heat transfer from the surface of fuel compacts to surrounding components such as graphite sleeves. The average moderator temperature rises slowly compared to the fuel temperature and the temperature rises about 4 C at 240 s after the test starts. Fig. 4.41 shows the transient of reactivity. The positive reactivity of 3.4 3 1024 k/k (5.3 cents) is inserted in 6.6 s by withdrawal of the central pair of control rods. However, the negative reactivity is inserted due to the Doppler effect with fuel temperature rise and the reactivity effect with moderator temperature rise

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Figure 4.41 Transient of the reactivity during control rods withdrawal test [24].

because of the reactor power increase. As shown in Fig. 4.41, the time when the positive reactivity by control rod withdrawal compensates with the negative one due to fuel and moderator temperature rise is about 140 s after the test starts. The reactor power gets to be almost stable at that time. The analytic result of reactor power shows a good agreement with the experiment one. The analysis model evaluates the reactor power about 1% higher than the experimental value for the control rods withdrawal test. There is a difference between analysis and experiment of reactor power just after the peak as shown in Fig. 4.39. According to the transient of total reactivity as shown in Fig. 4.41, the analysis result shows a time lag of about 10 s to the experiment result. It is thought that the fuel temperature calculation causes the differences of the reactor power between analysis and experiment because the total reactivity transient just after the reactor power peak is subject to the fuel temperature transient as shown in the Fig. 4.41. More accurate simulation of reactivity insertion events can be expected by an improvement of fuel temperature calculation model for the fuel compact. The reactivity insertion test-control rod withdrawal test revealed that the increase of reactor power by withdrawing the control rods is restrained by only the negative feedback of the core without operating the reactor power control systems. Then the temperature transition of core components is slow due to their large heat capacity.

4.5.6 Coolant flow reduction test—gas circulators trip test 4.5.6.1 Objective and test procedure The gas circulators trip test demonstrates that a rapid decrease of the coolant flow rate brings the reactor power to a stable level without a reactor shutdown and the transition of fuel temperatures is slow. Obtained test data are used for the development and validation of codes for safety evaluation of HTGRs.

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High Temperature Gas-cooled Reactors

In the gas circulators trip test, coolant flow rate is reduced by running down one and two out of three gas circulators with the reactor power control system being blocked. In this test, scram set values of primary coolant flow rate of PPWC, core differential pressure, and reactor outlet coolant temperature in case the primary coolant flow decreases are modified to prevent an automatic start of the reactor scram in the course of the tests. At the same time, power supply for driving all the control rods is cut off to prevent withdrawal of control rods by mistake. The modification of scram set values and cut-off of power supply driving the control rods are completed automatically upon selection of the test mode switch. Major measurement points and instrumentation are shown in Fig. 4.38.

4.5.6.2 Test results Test results of the coolant flow reduction test by tripping gas circulators in the reactor power of 9 MW are compared with analysis results as shown in Fig. 4.42 [24]. The flow rate of the primary circuit decreases immediately after tripping of gas circulators. Ten minutes elapsed; the reactor power is diminished to a stable level from the initial power of 9 MW due to the negative feedback effect increasing the reactor core temperature along with decreasing of the coolant flow rate. The reactor scram is not activated during the test. The analysis result of reactor power showed a good agreement with the experiment one. The coolant flow reduction test by the gas circulators trip revealed that the decrease of coolant flow rate brings the reactor power to a stable level by negative feedback of the core against the temperature arise without reactor shutdown and other control system, and the temperature transition of core components is slow due to their large heat capacity.

Figure 4.42 Transient of the reactor power and the primary coolant flow during gas circulators trip test [24].

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4.5.7 Loss of forced cooling test 4.5.7.1 Objective and test procedure The loss of forced cooling test is aiming to demonstrate that the reactor power decreases and gets to stable by the negative reactivity feedback effect, and the transient of the fuel temperature is slow under the condition of all-three-gas-circulator trip and the loss of forced cooling in the core without reactor scram. The test also demonstrates that the reactor power decreases immediately against its brief increase, which is posed by the positive reactivity insertion along with decrease of the fuel and moderator temperatures, and Xenon concentration, by the negative feedback effect. In the loss of forced cooling test, all three helium gas circulators of the HTTR are tripped by a signal from the central control room while deactivating all reactor power control systems in order to prevent an abnormal reduction in the primary coolant flow rate from inducing reactor scram. The loss of forced cooling test consists of tests with and without the active function of the VCS. The test simulating the loss of forced cooling with 1 out of 2 systems of the VCS inactive has so far been conducted. A test will be conducted with the deactivation of both active cooling trains of the VCS, and therefore, it tests one of the serious accidents typically included in the safety evaluation of HTGR, namely, passive decay heat removal in the case of a blackout of all cooling systems. The loss of forced cooling test is performed in increments of initial reactor thermal powers: the test case of 9 MW has already been performed as Run 1; and the test case of 30 MW will be performed as Run 2.

4.5.7.2 Test results In the loss of forced cooling with the active function of the VCS, the reactor power decreases immediately by the negative feedback of the Doppler reactivity posed by increase of the fuel temperature after the coolant flow rate gets to approximately zero through the loss of forced coolant function as shown in Fig. 4.43 [5]. Here, the reactor power includes only the reactor power without decay heat (neutron reactor power). The reactor power approaches to zero by the negative reactivity in the whole core due to reduction of reactivity by the moderator temperature and the Xenon concentration while the Doppler reactivity is positive inserted by decrease of the fuel temperature. Afterward, the reactor power increases slowly 4 h later after the test starts and the reactor power reaches at about 1% 8 h later. The inner surface temperature of the permanent reflector located at the side of the hot plenum increases after the test starts and decreases slowly 2 h later. The surface temperature of the RPV decreases along with the time at any position measured. This result shows that the heat generated inside the core is transferred by radiation from the RPV to the VCS. In the loss of forced cooling with the inactive function of 1 out of 2 systems of the VCS, the reactor response is similar to that with the active function of the VCS because the VCS is design to maintain its cooling performance needed with 1

310

High Temperature Gas-cooled Reactors

Figure 4.43 Analytical and experimental results of the reactor power transient in the LOFC test at 9 MW [25].

system. The temperature of cooling water in the inactive VCS increased about 10 C, which is sufficiently lower than the design temperature of the equipment. The loss of forced cooling test revealed that the decay heat generated in the core could be absorbed into the large heat capacity, transferred through the graphite core and the RPV, emitted by thermal radiation from its outer surface and removed to the active VCS; therefore the core at 9 MW is never exposed to the danger of a core melt, and the reactor power is stabilized spontaneously. The test with 1 out of 2 VCS inactive revealed that the VCS can remove the radiation heat from the RPV with 1 system as designed.

References [1] Y. Tachibana, et al., Procedure to prevent temperature rise of primary upper shielding in high temperature engineering test reactor (HTTR), Nucl. Eng. Des. 201 (2000) 227 238. [2] N. Fujimoto, et al., Experience of HTTR construction and operation-unexpected incidents, Nucl. Eng. Des. 233 (2004) 273 281. [3] Y. Miyamoto, et al., Demonstration Tests for HTGR Fuel Elements and Core Components with Test Sections in HENDL, Japan Atomic Energy Research Institute, JAERI 1333, 1995. [4] N. Nojiri, et al., Characteristic test of initial HTTR core, Nucl. Eng. Des. 233 (2004) 283 290. [5] K. Yamashita, et al., Nuclear design of the High-Temperature Engineering Test Reactor (HTTR), Nucl. Sci. Eng. 122 (1996) 212 228. [6] K. Yamashita, et al., Startup core physics tests of High Temperature Engineering Test Reactor (HTTR), (I) test plan, fuel loading and nuclear characteristics tests, J. Atom. Energy Soc. Jpn. 42 (1) (2000) 30 42.

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[7] N. Fujimoto, et al., Startup core physics tests of High Temperature Engineering Test Reactor (HTTR), (I) test plan, fuel loading and nuclear characteristics tests, J. Atom. Energy Soc. Jpn. 42 (5) (2000) 458 464. [8] IAEA Technical Document, Evaluation of High Temperature Gas Cooled Reactor Performance: Benchmark Analyses Related to Initial Testing of the HTTR and HTR10, IAEA-TECDOC-1382, 2002. [9] T. Mori, et al., Vectorization of continuous energy Monte Carlo Method for neutron transport calculation, J. Nucl. Sci. Technol. 29 (4) (1992) 325 336. [10] K. Kunitomi, et al., Development of new type of HTGR, in: The 73rd JSME Fall Annual Meeting, 1995. [11] S. Nakagawa, et al., Performance test of HTTR, Nucl. Eng. Des. 233 (2004) 291 300. [12] Y. Tachibana, et al., Plan for first phase of safety demonstration tests of the high temperature engineering test reactor (HTTR), Nucl. Eng. Des. 224 (2003) 179 197. [13] S. Hanawa, et al., Experimental and analytical study on thermal displacement applied to floating support unit, in: Proceedings of the 8th International Conference on Nuclear Engineering (ICONE-8), Baltimore, MA, 2000. [14] K. Sawa, et al., Fabrication of the first-loading fuel of the high temperature engineering test reactor, J. Nucl. Sci. Technol. 36 (1999) 683 690. [15] Y. Tachibana, et al., Test Plan Using the HTTR for Commercialization of GTHTR300C, JAEA-Technology 2009-063, 2009. [16] D. Tochio, et al., Result of Long-term Operation of HTTR-Rated/Parallel-Loaded 30days Operation, JAEA Technology 2009-005, 2009. [17] M. Goto, et al., Long-term high-temperature operation of the HTTR, Nucl. Eng. Des. 251 (2012) 181 190. [18] IAEA, Fuel Performance and Fission Gas Behavior in Gas Cooled Reactors, IAEA TECDOC-978, 1997. [19] S. Fujikawa, et al., Achievement of reactor outlet coolant temperature of 950 C in HTTR, J. Nucl. Sci. Technol. 41 (2004) 1245 1254. [20] S. Saito, et al., Design of High Temperature Engineering Test Reactor (HTTR), JAERI 1332, 1994. [21] K. Okumura, et al., SRAC2006: A Comprehensive Neutronics Calculation Code System, JAEA-Data/Code 2007-004, 2007. [22] Nuclear Safety Commission, Guidelines for Safety Evaluation of LWR Power Plants, 1989. [23] Nuclear Safety Commission, Guidelines for Reactor Siting Evaluation, 1989. [24] S. Nakagawa, et al., Safety demonstration tests using high temperature engineering test reactor, Nucl. Eng. Des. 233 (2004) 301 308. [25] K. Takamatsu, et al., Experiments and validation analyses of HTTR on loss of forced cooling under 30% reactor power, J. Nucl. Sci. Technol. 51 (2014) 1427 1443. [26] T. Takeda, et al., Development of Analytical Code “ACCORD” for Incore And Plant Dynamics of High Temperature Gas-Cooled Reactor, Japan Atomic Energy Research Institute, JAERI-Data/Code 96-032, 1996.

R&D on commercial high temperature gas-cooled reactor

5

Jun Aihara1, Takeshi Aoki1, Yuji Fukaya1, Minoru Goto1, Yoshiyuki Imai1, Yoshitomo Inaba1, Yoshiyuki Inagaki1, Tatsuo Iyoku1, Yu Kamiji1, Seiji Kasahara1, Shinji Kubo1, Kazuhiko Kunitomi1, Naoki Mizuta1, Odtsetseg Myagmarjav1, Tetsuo Nishihara1, Hiroki Noguchi1, Hirofumi Ohashi1, Nariaki Sakaba1, Koei Sasaki1, Hiroyuki Sato1, Taiju Shibata1, Junya Sumita1, Yukio Tachibana1, Shoji Takada1, Tetsuaki Takeda2, Hiroaki Takegami1, Nobuyuki Tanaka1, Shohei Ueta1 and Xing Yan1 1 Sector of Fast Reactor and Advanced Reactor Research and Development, Japan Atomic Energy Agency, Ibaraki, Japan, 2Graduate Faculty of Interdisciplinary Research, Research Faculty of Engineering, Department of Mechanical Engineering, University of Yamanashi, Yamanashi, Japan

High temperature gas-cooled reactor (HTGR) can be used in various ways such as hydrogen production and process heat supply as well as power generation. For the commercial HTGR, R&D on the system and safety designs, key components for heat application and advanced fuel have been carried out by JAEA as shown in the following items: 1. System design for power generation Conceptual designs of a steam cycler plant named HTR50S and a gas turbine power plant named GTHTR300 with are performed. The HTR50S has a thermal power of 50 MW and a reactor outlet coolant temperature of 750 C. The GTHTR300 a thermal power of 600 MW and reactor outlet coolant temperatures of 850 and 950 C. 2. System design for cogeneration Conceptual cogeneration designs of hydrogen production by the thermochemical water splitting process, power generation by the gas turbine, and seawater desalination by the multistage flash (MSF) process are performed. Furthermore, a conceptual design of a renewable hybrid system coupling the gas turbine and the solar power generations is described. 3. System design for steelmaking A conceptual design of the nuclear hydrogen steelmaking (NHS) process is performed. In the process, hydrogen produced by the thermochemical water splitting process is used as a reducing agent for steelmaking. 4. Safety design for connection of HTGR and heat application system The safety standards such as the safety requirements, the acceptance criteria, etc. towards the commercialization of HTGR is described. Furthermore, the high temperature engineering test reactor (HTTR) cogeneration demonstration is planned to establish the safety standards. 5. Gas turbine technology for power generation The basic technologies of key components such as a helium gas compressor and a control system of a magnet bearing have been High Temperature Gas-cooled Reactors. DOI: https://doi.org/10.1016/B978-0-12-821031-4.00005-1 © 2021 Elsevier Inc. All rights reserved.

314

6.

7.

8.

9.

10.

11.

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developed. By the experimental and numerical studies with a 1/3-scale and four-stage test model, main specifications of the helium gas compressor were decided. By numerical study, it was confirmed that a multiinput multioutput (MIMO) controller is suitable for the control system of the magnet bearing. Iodinesulfur (IS) process technology for hydrogen production Key components made of industrial materials such as a sulfuric acid decomposer with a ceramic reaction tube was developed. A hydrogen production test facility made of industrial materials was fabricated, and continuous hydrogen production was performed for 31 hours at a hydrogen production rate of 20 NL/h. For improvement of hydrogen production efficiency, R&D on an electrodialysis stack and a hydrogen permselective membrane have been performed. System integration technology for connection of heat application system and HTGR Four key technologies for the safe and economical connection have been developed as follows: (1) a control method to mitigate the temperature fluctuation in the helium gas coolant caused by chemical reactors, (2) estimation of tritium permeation from HTGR to a hydrogen production system, (3) a countermeasure against explosion of combustible gas, and (4) a high temperature isolation valve (HTIV) to separate reactor and hydrogen production system in accidents. Prevention technology for air ingress during depressurization accident The SCAD system for the GHTR300C system was proposed by JAEA. When not only the localized natural convection but also natural convection was generated, the onset time of natural circulation becomes short. Therefore, it is important to know that the localized natural convection or circulation is generated or not. Advanced fuel technology for high burnup R&D on the fuel to increase economy of HTGRs have been performed to maintain their integrities at the burnup of three to four times higher than that of the conventional HTTR fuel (33 GWd/t in maximum). In the design of the fuel, the thickness of the coating layer was examined in order to achieve a high burnup based on the irradiation test so far. Advanced fuel technology for plutonium burner The concept based on HTGR technology, “Clean Burn”, is proposed by JAEA. The Clean Burn concept can use surplus plutonium as a fuel without mixing it with uranium matrix by employing an inert matrix fuel (IMF) and a tightly coupled fuel reprocessing and fabrication plants. Thus, surplus plutonium alone will be incinerated effectively, while generation of plutonium from the uranium matrix is avoided. Advanced fuel technology for reduction of high-level radioactive waste (HLW) The effective waste loading method for direct disposal is proposed. By taking into account features such as high burnup, high thermal efficiency, etc., the number of HLW canister generations and its repository footprint are evaluated by burnup fuel composition, thermal calculation, and criticality calculation in repository. As a result, it is found that the number of canisters and its repository footprint per electricity generation can be reduced by 60% compared with the light water reactor (LWR) representative case for direct disposal.

5.1

System design for power generation

5.1.1 Introduction The use of ceramic-coated fuel particle, graphite moderator, and helium coolant offers high temperature heat supply capability to HTGR. The capability provides

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higher power conversion efficiency. A number of HTGR power plants with steam and gas turbines have been designed. JAEA conducted a conceptual design of an HTGR steam cycle power plant named HTR50S from 2010 to 2012 aiming to deploy in developing countries in 2030s [1]. The design is emphasized on using the existing technology and experience obtained in the HTTR construction and operation to minimize initial construction uncertainty and potential delay. The employment of gas turbine power conversion system has several advantages over the use of steam turbine power generation system in HTGR. First, the system is the most efficient and economical power generation possible by HTGR. Application of direct Brayton cycle gas turbine enables 50% efficiency at least count of components. Second, wide range of heat application is possible without drawbacks in efficiency. Third, the employment can exclude water-related safety events, that is, water ingress to reactor core, from HTGR. After a several design studies for gas turbine high temperature reactor were conducted in JAERI over a period of several years, the reference design of a nominal 300 MWe gas turbine high temperature reactor system named GTHTR300 had been performed by JAEA together with major domestic industries from 1998 to 2008 aiming to complete the basic design for domestic deployment of the lead plant in coming decades [2]. The design is based on comprehensive experience and know-how in reactor design, construction, operation, and maintenance, which had been acquired through research and development in the JAEA’s HTTR project.

5.1.2 HTR50S: HTGR steam cycle power plant Fig. 5.1 shows the system configuration of the HTR50S including the reactor, steam generator (SG), steam cycle power conversion system, stop cooling system, and vessel cooling system (VCS). Table 5.1 lists major specifications of the HTR50S. The HTR50S operates at reactor thermal power of 50 MW with reactor outlet temperature of 750 C. The reactor outlet temperature is determined to allow the application of Alloy 800H to tube material for a SG. The reactor inlet temperature of 325 C is also selected to enable the use of SA533/SA508 for the reactor pressure vessel (RPV) material. The heat generated in the core is transferred to the steam cycle power conversion system through the SG. A regenerative cycle with a threestage extraction turbine is used. A reheating cycle, which is used in the previous HTGR power conversion systems, for example, the Thorium Hochtemperatur Reaktor and the Fort St. Vrain reactor, is effective for increasing power generation efficiency, however, the cycle is not employed to avoid complexity of the SG configuration, increase in the number of penetrations in reactor containment vessel. Fig. 5.2 shows cross-sectional views of the reactor. The reactor core consists of six layers of hexagonal blocks to facilitate fuel shuffling. The core is configured cylindrically with 30 fuel columns, 18 replaceable reflector columns, and 13 control rod guide columns. Each fuel column has 33 fuel channels with fuel elements and coolant holes forming circular flow paths. TRISO-coated fuel particles with UO2 fuel kernels are formed into fuel compacts and inserted into graphite sleeves.

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533°C 12.0 MPa 69.6 kg/s

Reactor containment Reactor Vessel cooling system

750°C 4.0 MPa 80.5 Steam kg/s generator

CS

17MWe

G

538°C 12.5 MPa 69.6 kg/s

325°C 4.04 MPa 80.5 kg/s

CS

Steam turbine

Generator

Condenser Deaerator

Stop cooling system

Steam dump system

Feed water heater

Feed water pump

60°C 0.95 MPa 54.1 kg/s

200°C 13.4 MPa 69.6 kg/s

Figure 5.1 System configuration of HTR50S. Table 5.1 Major specifications of HTR50S. Reactor power Reactor inlet temperature Reactor outlet temperature Reactor outlet coolant pressure Coolant flow rate Average core power density Fuel element Uranium enrichments Number of enrichments Average fuel burnup Refueling interval Power generation Power generation efficiency Turbine inlet temperature Turbine inlet steam pressure

50 MWt 325 C 750 C 4.1 MPa 22.3 kg/s 3.5 W/m3 Sleeved 4.3, 6.6, 9.4 wt.% 3 33 GWd/t 2 years 17.2 MWe 34 533 12.0 MPa

Control rods and reserve shut down system channels are located in the reflector blocks on the inner and outer boundaries of the fuel region. Permanent side reflectors surround the outside of outer replaceable reflectors. Burnable poisons (BPs) are installed below the dowel pins, which are allocated on the corners of the fuel block. The reactor coolant enters the RPV through the outer side of a coaxial hot gas duct, flows upward through the flow path between the permanent side reflectors and side shielding blocks, and reverses the direction at the upper plenum. The coolant temperature increases as it flows downward through the coolant channels, joins at the lower plenum, and exits the reactor through the inner side of the coaxial hot gas

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Stand pipe Shroud

RPV

317

Fuel block

Permanent reflector

Reactor core Coaxial hot gas duct

Replaceable Control rod reflector guide block

Figure 5.2 Cross-sectional views of HTR50S reactor.

duct. The upper shroud is installed to prevent the contact of hot gas with the RPV during the pressurized loss-of-forced circulation events. Fig. 5.3 shows cross-sectional views of the SG. Table 5.2 shows major specifications. The SG is a helically coiled counter-type heat exchanger. The primary helium gas flows in the shell side and the secondary water/steam inside the tubes. Hightemperature helium gas from the reactor is introduced to inner side of the inner shroud and transfers heat to secondary water/steam through helically coiled tubes. The helium gas flows out from the heat transfer section and is guided to helium gas circulators and flows back to the SG. The pressurized helium gas is introduced to the flow path between the outer shell and inner shroud and flows back to the reactor through inner piping in the coaxial hot gas duct. Helium gas temperature at the outlet of circulator is 325 C, which is sufficiently low to maintain the outer shell temperature to use MnMo steel. The tube material used is STBA24 for the boiling section and Alloy800H for super heating section. A bimetallic welding is used to connect these tubes. Fig. 5.4 shows a system layout of the VCS. The VCS is one of the engineered safety systems in the HTR50S designed to remove decay and residual heat from the reactor core to maintain the integrity of RPV, reactor internal structure, as well as prevent undue failure in fuel. In addition, the VCS operates not only in accident but also during normal operation to maintain concrete temperature of reactor silo below the allowable temperature. The VCS consists of cooling panels, natural draft air coolers, containment isolation valves, etc. The system is fully passive and uses natural circulation of water to remove heat from cooling panels, which receives heat from the RPV via thermal radiation and natural convection. The single failure criterion is applied to the system because the natural circulation flow will be lost in case of active containment isolation valves. Hence, the VCS consists of two independent

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High Temperature Gas-cooled Reactors

From reactor 750°C, 4 MPa Steam 538°C, 12.5 MPa 69.6 t/hr Tube; Superheating section Tube; Boiling section

From circulator

To circulator

From circulator

To circulator

Feed water 200°C, 13.3 MPa

Figure 5.3 Cross-sectional view of HTR50S steam generator.

Table 5.2 Major specifications of HTR50S steam generator. Operating conditions Primary helium gas flow rate Secondary steam/water flow rate Primary helium gas inlet temperature Primary helium gas outlet temperature Primary water inlet temperature Secondary steam outlet temperature Steam outlet pressure Heat capacity

22.4 kg/s 19.3 kg/s 750 C 325 C 200 C 538 C 12.5 MPa 50 MW

Design specifications Tube outer diameter Tube thickness Tube number Inner shell inner diameter Inner shell outer diameter

38.8 mm 3.5 mm 36 0.664 m 1.436 m

Material Tube Outer shell

STBA24/Alloy800H SCMV4-2

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Natural draft cooling tower

Natural draft cooling tower Containment vessel

Natural circulation of water

Natural circulation of water

Cooling panels

Figure 5.4 System layout of HTR50S vessel cooling system.

systems. The natural draft air coolers are installed on the roof of reactor building to provide sufficient head for the natural circulation. The system is designed to maintain the working fluid in single phase and the water pressure is maintained below 1.0 MPa by pressurizers.

5.1.3 GTHTR300: HTGR gas turbine power plant Fig. 5.5 shows the plant layout of GTHTR300 consisting of the reactor, gas turbine generator (GTG) module, and heat exchanger (HTX) module. Table 5.3 lists major specifications of the GTHTR300. The gas turbine consists of an axial-flow turbine and an axial-flow compressor. It drives an electric generator on a common shaft. The HTX module consists of heat exchanger vessel, offset fin-type recuperator, and precooler with low finned heat transfer tubes. Because the power conversion cycle rejects waste heat over a wide temperature range at the precooler, district heating or desalination can be carried out with the use of the waste heat. Fig. 5.6 shows cross-sectional views of the reactor [3]. The core consists of eight layers of hexagonal fuel blocks made of graphite and is installed between the replaceable reflector blocks. Fuel blocks are arranged in an annular region and neutron reflector blocks are arranged inside and outside of the fuel region. The coated fuel particles with UO2 fuel kernels are formed into fuel compacts and formed as monolithic fuel rods. The reactor coolant is introduced to annular flow channels between fuel rods and coolant channels devised in fuel blocks. Channels are provided in the reflector blocks for insertion of control rods and B4C pellets, which are used to shut down reactor in case of control rod failure. Permanent side reflectors are installed outside of outer replaceable reflectors. BPs are placed under dowel pins, which are allocated on the corners of the fuel block.

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High Temperature Gas-cooled Reactors

Reactor

HTX module Reactor core Recuperator

GTG module

Turbine

Precooler

Compressor

Generator

Figure 5.5 Plant layout of GTHTR300 [2].

Table 5.3 Major specifications of GTHTR300.

Reactor power (MWt) Reactor inlet temperature ( C) Reactor outlet temperature ( C) Reactor inlet coolant pressure (MPa) Coolant flow rate (kg/s) Average core power density (W/m3) Fuel element Uranium enrichment (wt.%) Number of enrichments Average fuel burnup (GWd/t) Refueling interval (years) Power generation (MWe) Power generation net efficiency (%)

Baseline design

High performance design

600 587 850 6.9 403 5.4 Monolithic 14.0 1 120 2 274 45.6

600 666 950 6.4 439 5.4 Monolithic 7.016.8 8 120 1.5 302 50.4

In order to use conventional steel material for the RPV, an innovative cooling method utilizing the unique characteristics of power conversion cycle is employed. The reactor coolant flows into the reactor at 587 C and is introduced to flow channels installed in permanent reflectors to avoid direct contact with the RPV. Another annulus flow path is devised between the RPV and core barrel. The low temperature, high pressure helium coolant discharged from the compressor is introduced to the flow path as

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Stand pipe Upper plenum shroud

Control rod insertion hole

321

Reserve shut down channel

Control rod guide block Fuel block

Upper plenum Upper shield Upper reflector Core barrel Coolant channel through permanent side reflector RPV

Lower reflector

Primary loop

Primary loop

Lower plenum Carbon block Core support plate

(A)

Inner replaceable reflector

Permanent side reflector

Outer replaceable reflector

(B)

Figure 5.6 Cross-sectional views of GTHTR300 reactor [3]. (A) Vertical cross-section. (B) Horizontal cross-section.

shown in Fig. 5.7. The coolant flows upward through the annulus and cools the RPV, core barrel, and control rod drive mechanisms. The flow is guided to inside of the control rod guide tube and joins up with main stream of reactor coolant. A small portion of the coolant from the compressor is introduced to the bottom of the reactor and used to maintain temperature of bottom metallic supports, for example, core support plate. These flows are driven by intrinsic pressure difference due to the cycle characteristics and do not require additional devices. The actual operating temperature of the RPV depends largely on the rate of bypassed cooling flow. The reference design selects a 0.2% bypassed cooling flow to keep the vessel operating temperature well below the vessel material design limit for conventional steel SA533/SA508 over the entire potential range of power operations. Fig. 5.8 shows the cycle diagram of the GTHTR300. The power conversion system of GTHTR300 employs closed, nonintercooled cycle based on the comparative cost performance evaluation for cycle configuration. Helium coolant heated at the reactor core flows through outside of the coaxial hot gas duct and provides heat to the turbine. The coolant expanded by the turbine flows through the recuperator and precooler, and flows into the compressor. A part of the compressed coolant is guided to a flow path inside the RPV in order to cool the RPV, while the remaining compressed coolant is heated at the recuperator and flows back to the reactor. Rejected waste heat of Brayton cycle at the precooler is used for district heating or desalinations. The cycle pressure ratio of 2.0, which can achieve optimum cost performance, is selected. The net plant efficiency of 45.6% is obtained with reactor outlet temperature of 850 C. The efficiency can be increased to 50.4% by updating key parameters of the cycle [4].

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High Temperature Gas-cooled Reactors

Figure 5.7 RPV cooling flow driven entirely by intrinsic pressure gradients [2].

First, reactor outlet temperature, namely, turbine inlet temperature can be increased to 950 C due to two major technical achievements: HTTR continuous reactor operation with reactor outlet temperature of 950 C [5] and commercial success of single-crystal Ni-based alloy used as turbine blade material in aircraft engine [6]. Second, the compressor efficiency is updated based on the result of the recently concluded development program in JAEA [7]. The update contributes to the gain of efficiency by 0.5%. A similar efficiency increase is expected in the turbine, which leads to efficiency increase by 0.4%. Finally, the recuperator effectiveness is optimized to obtain 0.9% increase efficiency with only a 10% increase in the recuperator size. Fig. 5.9 shows the bird’s eye view of the GTHTR300 gas turbine. The gas turbine employs single-shaft, axial flow design with six-stage turbine and twenty-stage compressor. Horizontal orientation is selected and the generator is connected in the rotating shaft via diaphragm coupling in order to reduce bearing loads. The generator is submerged in a pressure vessel to avoid shaft penetration into the vessel with an aim to reduce helium gas leakage. Instead, lubricant-free bearings are installed for the shaft support. Magnetic bearings were selected for the baseline design and the control system was tested in a rotor test rig.

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323

&RROLQJIORZIRUWXUELQH Electric generator Compressor

Turbine

&RROLQJZDWHUIORZ

Precooler Reactor &RROLQJIORZIRUUSY

Recuperator

Figure 5.8 GTHTR300 cycle diagram.

Diaphragm coupling Compressor Journal bearing

Generator Thrust bearing Journal bearing Turbine

From reactor To reactor

From precooler To recuperator

Figure 5.9 Schematic view of GTHTR300 gas turbine.

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High Temperature Gas-cooled Reactors

Either vertical or horizontal orientation can be selected for the turbomachine. Past direct cycle modular HTGR designs preferred vertical orientation. The turbomachinery orientation has impact on rotor dynamics, bearing requirements, maintainability, and plant building arrangement. Optimum building arrangements were studied for the horizontal and vertical installation options, and major difference between the arrangements was limited to the turbine building area. Taking account of the spaces required for both installation and maintenance, the horizontal design requires a larger plant area in turbine building. On the other hand, the vertical design needs a taller turbine building elevation. The volume of overall plant building for the vertical design is about 6% smaller than the horizontal design, and the vertical design can potentially save 1% in plant construction cost. Nonetheless, the vertical turbomachinery installation has significant disadvantage in bearing design because the design poses undue thrust loads in magnetic bearing and auxiliary catch bearing. In case of horizontal arrangement, static and dynamic loads are not jointly concentrated on any one bearing. Hence, the loads on bearing are reduced to within the design limits of the present technologies. In addition, the horizontal turbomachinery design has large advantage due to the considerable experience related to installation, operations, and maintenance of machines in industry. In conclusion, the horizontal orientation was preferred in GTHTR300. Overall cycle efficiency decreases rapidly as the turbine cooling flow increases. Hence, a turbine cooling scheme (see Fig. 5.10), which can limit the amount of flows in turbine rotor and casing to less than 1% of main flow, is employed for the present turbine design. The required amount of flow for turbine cooling is low because the turbine only has six stages due to the low expansion ratio. In addition, externals of turbine casing including inlet and exhaust pieces are cooled by main flow at about 136 C that is discharged from the compressor and therefore such dedicated flow is only needed for internals of casing. The casing internals such as disc and bearing are cooled with flow stripped from the main flow. The recovered leak flow from the turbine shaft is used to cool the exhaust scroll. The analysis by a finite element method confirmed that the temperatures, related stress, and displacement of the turbine casing are below design criteria.

Figure 5.10 Details of turbine cooling flowpath designed to minimize cooling flow requirements for disks, casing, and bearing [2].

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325

Table 5.4 describes major specifications of the turbine. Turbine design employs efficient reaction stages with similar tip speed and stage loading based on conventional designs. As a result, a six-stage turbine with uncooled blades made of conventional directionally solidified alloy is employed. The turbine design also employs advanced design techniques in order to improve the performance. A three-dimensional blade design is used to optimize secondary flows whose magnitude of loss is in the same order as that of the profile loss in case of conventional blade design. The stator uses bowed blades of high aspect ratio to reduce end wall losses significantly while the rotor incorporates blades with increased inlet flow angles toward the hub to mitigate the adverse Mach number effect in that region. As shown in Fig. 5.11, the secondary loss in the present helium gas turbine only Table 5.4 Major specifications of GTHTR300 turbine. Rotational speed Stage number Rotor diameter (1st stage/6th stage) Rotor blade outer diameter (1st stage/6th stage) Boss ratio (1st stage/6th stage) Rotor blade height (1st stage/6th stage) Stator blade height (1st stage/6th stage) Number of blades (stator/rotor) Flow coefficient (1st stage/6th stage) Load coefficient (1st stage/6th stage)

3600 rpm 6 1844/1750 mm 156/250 0.855/0.778 mm 156/250 mm 150/240 mm 82/80 04.43/0.46 1.52/1.16

Figure 5.11 Comparison of stage-wise aerodynamic losses of helium gas turbines employing conventional and advanced blades, indicating potential of incorporating advanced blading techniques to reduce secondary losses [2].

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High Temperature Gas-cooled Reactors

Table 5.5 Major specifications of GTHTR300 compressor. Rotational speed Stage number Rotor diameter (1st stage/20th stage) Rotor blade outer diameter (1st stage/20th stage) Boss ratio (1st stage/20th stage) Rotor blade height (1st stage/20th stage) Stator blade height (1st stage/20th stage) Number of rotor blades (1st stage/20th stage) Number of stator blades (1st stage/20th stage) Flow coefficient Work coefficient Reaction

3600 rpm 20 1500/1500 mm 1704/1645 0.88/0.91 mm 102/72.5 mm 101/71.5 mm 72/90 94/116 0.51 0.63 0.75

contributes to a small fraction of the total loss in all six stages. The stage efficiency is improved by a 0.5%1.0% from the performance of conventional design. Table 5.5 lists the major specifications of compressor. Although design parameters of helium gas compressor are similar to those of the industrial gas turbine compressors, the compressor design has some unique challenges. For example, the flow path of helium gas compressor is required to be parallel with high hub-to-tip ratios (around 0.9) throughout the twenty stages, whereas typical compressor only needs a similar flow path in the rear stages. The characteristics pose complexity for the consideration of secondary flow in the design to meet requirement of the efficiency and stall margin. First, a high reaction airfoil is used to achieve reduction in flow path friction loss and sufficient stall margin as shown in Fig. 5.12A. Second, the advanced blade design is incorporated including the blade-end bends to eliminate flow separation and the blade over-camber to compensate for flow distortion near the end wall (see Fig. 5.12B and C). Third, blade row solidity and aspect ratio are optimized for efficiency and flow stability. Finally, inlet and outlet geometries are optimized by using CFD to control pressure losses in these locations (Fig. 5.12D). The HTX module contains the recuperator, precooler, and control valves in a vertical steel pressure vessel as shown in Fig. 5.13. Offset-strip fin heat exchanger is selected for the recuperator to achieve recuperator effectiveness of 95% and minimize pressure loss as well as size of heat exchanger modules. The details of heat exchanger modules are also shown in Fig. 5.13. The recuperator is made up of six heat exchanger modules with an 8 m height that operates in parallel. The heat exchanger modules use a fin with a size of 1.2 mm height and 1.2 mm pitch. The fins are blazed with plates with 0.5 mm thickness. The precooler is based on a helically coiled tube bundle with water in the tubes and helium gas in the shell. A key design benefit in the nonintercooled cycle is that the cycle rejects waste heat over a broad high temperature range. This increases effective logarithmic mean temperature difference and decreases surface area requirement for the precooler. Low-finned tube is employed for the heat transfer

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327

Figure 5.12 Elements of high-performance helium gas compressor design approach [2]. (A) Reduction of friction loss using high reaction air-foil. (B) Blade-end bends for elimination of flow separation. (C) Blade over-camber for compensation of flow distortion near end wall. (D) Inlet and outlet geometries.

tubes and the volume and weight are reduced by 20%30% compared to the smooth tube design variant. Fig. 5.14 shows the cross-sectional view of the GTHTR300 VCS. The VCS is designed to remove heat dissipated at outer surface of the RPV by using cooling panels installed inside the reactor silo. The fully passive air-cooled system is

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High Temperature Gas-cooled Reactors

970 mm Control valves

0.15 mm 1.2 mm

Recuperator 0.5 mm

From compressor

Heat transfer module

Fin detail

To reactor

Precooler Pressurized water

To compressor From turbine

Arrangement of heat transfer modules

Figure 5.13 Cross-sectional view of GTHTR300 heat exchanger module.

employed by utilizing natural circulation of air from outside of the reactor building. The air introduced to the system flows downward through the flow path between low temperature panel and reactor silo and reverses the direction at the bottom. The air forms upward flow due to the heat received by high temperature panel and flows out to the atmosphere though ducts. The flow paths are formed in rectangular shape to assure integrity against load induced by depressurization accident. Fig. 5.15 shows the comparison of plant construction cost, fuel cost, and power generation cost between GTHTR300 baseline design and a pressurized water reactor (PWR) of 1300 MWe [8]. The cost estimation assumes a plant life of 40 years, the discount rate of 3%, and load factor of 80%. The capital cost is 1.57 Yen/kWh, which is 25% lower than that of PWR (2.13 Yen/KWh) largely due to the high power generation efficiency. The operation and maintenance (O&M) cost of GTHTR300 are 35% lower (1.11 Yen/ KWh) than those of PWR (1.71 Yen/KWh) because of the greater power generation efficiency and less count of components. The fuel cost of GTHTR300 is 1.46 Yen/KWh, which is equivalent to that of PWR (1.46 Yen/KWh). The cost of enrichment, fabrication, and reprocessing in GTHTR300 are higher of the two; however, the cost per kilowatt hour becomes comparable due to the higher efficiency in electricity generation. As a result, power generation cost of GTHTR300 is 20% lower (4.14 Yen/KWh) than that of PWR (5.3 Yen/KWh). As indicated in the result of the study, the impact of power generation efficiency on the cost is significant. The introduction of high performance design, a 50% power generation efficiency GTHTR300 with reactor outlet temperature of 950 C, can reduce further the power generation cost by 10% [9].

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329

Outlet

Inlet

Intake duct Exhaust duct Upper panel

Reactor cavity

Insulator

High temp. panel

Low temp. panel

Reactor silo

Conducon Lower panel

Radiaon Low temperature air flow High temperature air flow

Figure 5.14 Cross-sectional view of GTHTR300 VCS [3].

5.2

System design for cogeneration

5.2.1 Introduction HTGR is suitable for industrial cogeneration due to its inherently safe characteristics. Radionuclides can be retained within plant by autonomous reactor shutdown and core cooling without any equipment actions against loss of offsite power or failure of piping in primary system. The increase in outlet temperature up to around 1000 C offers a range of possibilities for industrial cogeneration High temperature heat can be used for highly efficient gas turbine power generation as described in the previous chapters. Large-scale hydrogen production by thermochemical water splitting process can be also carried out to make industrial uses such as steel making plant and transportation fuel. Process steam may be supplied in a wide range of temperature and pressure demanded by petroleum

330

High Temperature Gas-cooled Reactors

Reactor components GTHTR300

Power conversion system Auxiliary system Electric and control system

LWR(PWR)

Buildings 0

10

20

30

Construction cost (10,000 Yen/kWe) U purchase, conversion Enritchment Fabrication MOX Storage Reprocessing Waste disposal

GTHTR300

LWR(PWR) 0.0

0.5 1.0 1.5 Fuel cost (Yen/kWh)

2.0

GTHTR300

Capital cost O&M cost Fuel cost

LWR(PWR)

0.0

1.0

2.0

3.0

4.0

5.0

6.0

Power generation cost (Yen/kWh)

Figure 5.15 Comparison of costs between GTHTR300 and PWR [8].

and chemical industries. The waste heat in gas turbine power generation cycle is expected to be utilized in seawater desalination and district heating. The following sections provide examples of the HTGR cogeneration plant design.

5.2.2 Hydrogen cogeneration Based on the basic design of GTHTR300, an HTGR gas turbine power generation plant, a design series named the GTHTR300C [10] was developed by extending the capability of the plant for hydrogen cogeneration. The schematic of the GTHTR300C is shown in Fig. 5.16. The plant shares common equipment technologies and system configuration with the GTHTR300 with the exception of an

R&D on commercial high temperature gas-cooled reactor

HTX module

Annular

Reactor

331

GTHTR300C Hydrogen production plant

IHX module

Control valves

Recuperator

process heat

helical tube bundle

IHX

GTG turbine module Gas

Recuperator

Precooler

Precooler

Compressor

Generator

GTHTR300C rev.080403

Figure 5.16 Schematic view of GTHTR300C reactor system.

Table 5.6 Major specifications of GTHTR300C.

Reactor power (MWt) Hydrogen production plant heat supply rate (MWt) Reactor inlet temperature ( C) Reactor outlet temperature ( C) Reactor inlet coolant pressure (MPa) Coolant flow rate (kg/s) Refueling interval (years) Cycle pressure ratio Power generation efficiency (%) Power generation (MWe) Hydrogen production efficiency (%) Hydrogen production (m3/h)

GTHTR300C-I

GTHTR300C-II

600 170 594 950 5.1 322 1.5 2.0 47 202 50 c. 31,900

600 371 594 950 5.1 322 1.5 2.0 38 87 50 c. 69,600

intermediate heat exchanger (IHX). The IHX is installed between the reactor and the gas turbine and used to transfer required portion of reactor thermal power to process heat application including hydrogen production process or to other hightemperature process heat applications. Major specifications for the hydrogen cogeneration plant are listed in Table 5.6. The IHX and the gas turbine are arranged in series to effectively use the heat and reduce the size of the IHX by setting large logarithmic mean temperature difference between the primary and secondary helium gases. There are two variants, the GTHTR300C-I and the GTHTR300C-II, for the hydrogen cogeneration plant design depending on the amount of hydrogen production. The pressure of the primary

332

High Temperature Gas-cooled Reactors

coolant is reduced from 7 to 5 MPa for the both plants. The reduction is selected because baseline design of the GTHTR300 gas turbine can be directly applicable due to the similarity in aerodynamic and mechanical conditions. This scaling approach enables the technologies developed for the power conversion system of the power generation plant to be applicable to that of hydrogen cogeneration plant. Table 5.7 provides major specifications for both the IHX designs. Figs. 5.17 and 5.18 show the IHX designs for the GTHTR300C-I and the GTHTR300C-II. A helically coiled tube shell, and tube-type vertical heat exchanger are used for both IHXs based on the experience of design and construction of the HTTR IHX [10,11]. The high temperature helium gas heated in the reactor flows into the downside of the IHX, flows upward around the tube bundles, and transfers heat to secondary helium gas. The primary helium gas flows out from the IHX through inner piping of coaxial hot gas duct and is introduced to gas turbine. The pressurized helium gas from the compressor is introduced to the annulus flow path between inner and outer shells of the IHX with cooling the shells. The secondary helium gas from the process heat application flows into the IHX and is distributed to the helically coiled heat transfer tubes. The helium gas increases its temperature as the gas flows downward through the tubes. The heated helium gas changes the flow direction at the hot header and flows upward though the central pipe. The thermal insulators are utilized for the center pipe and inner shell in order to reduce the metallic temperature. The secondary helium gas flowing out from the IHX is delivered to the hydrogen production plant decreasing its temperature due to thermal dissipation of helium gas to atmosphere through the heat transport piping. In order to maximize the plant economics, the design of heat transport piping should consider not only the construction cost of heat transport piping but also the impact of heat loss on reduction in amount of hydrogen production. In order to develop an economical design for heat

Table 5.7 Major specifications of GTHTR300C IHX.

Heat exchange rate (MWt) Heat transfer area (m2) Working fluid (tube/shell) Flow rate (tube/shell, kg/s) Primary helium gas temperature (inlet/outlet,  C) Secondary helium gas temperature (inlet/outlet,  C) Primary inlet pressure (MPa) Secondary inlet pressure (MPa) Tube type Tube diameter (mm) Tube pitch (mm) Tube thickness (mm) Finn height (mm) Tube number

GTHTR300C-I

GTHTR300C-II

170 1448 He/He 80/324 950/850 500/900 5.0 5.15 Bare 45.0 65.0 5.0  724

370 3663 He/He 177/325 950/719 475/900 4.98 5.13 Finned 31.8 40.0 4.0 1.5 3685

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333

Secondary helium (to H2 plant)

Secondary helium (from H2 plant)

Φ1500

Primary helium (from compressor)

Inner shell

Primary helium (to turbine)

Thermal insulator

Heat transfer tube 22,000

Inner diameter Φ1840 Outer diameter Φ4570

Outer shell

Φ5500

Primary helium (to reactor)

Hot header Primary helium (from reactor)

Figure 5.17 Cross-sectional view of GTHTR300C-I IHX.

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High Temperature Gas-cooled Reactors

Secondary helium outlet Secondary helium inlet

Liner Tube sheet

Tube sheet support Center pipe Outer shell Primary helium inlet Inner shell Primary helium outlet Outer shell

Insulation

Primary helium outlet Baffle plate Primary helium inlet Insulation Heat transfer tube Hot header

Figure 5.18 Cross-sectional view of GTHTR300C-II IHX [11].

transport system for hydrogen production, comparative studies are conducted for different types of design for the heat transport piping. Fig. 5.19 provides the cross-sectional views of the heat transport piping considered in the study. Two types of structures, which are employed in the HTTR primary system, are selected. The first is a coaxial hot gas duct consisting of outer insulator, outside tube, inner tube, inner liner, and internal insulation. The hightemperature helium gas flows inside the liner made of Hastelloy XR. Internal insulation material made of a ceramic fiber composed of SiO2 and Al2O3 surrounds the inner tube. An outer tube is assigned to be the pressure boundary; however, its temperature is relatively low because high temperature helium gas is isolated by the inner insulation and the outer tube is cooled by low-temperature helium gas. As a result, SUS316 is used for the tube material.

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Inner tube

Liner

Inner insulator

335

Liner

Inner insulator

High temperature helium

High temperature helium

Low temperature helium Outer tube

Outer insulator

Outer tube

(A)

Outer insulator (B)

Figure 5.19 Cross-sectional views of heat transport piping. (A) Coaxial hot gas duct. (B) Inner insulated single piping.

3.0

Amount of material (×103 kg)

Heat loss (MW)

4.0

Single piping without outer insulator

2.0 Coaxial hot gas duct

1.0

Single piping with outer insulator

0.0 1

1.2

1.4

1.6

1.8

1200 Coaxial hot gas duct

Single piping with outer insulator

900

600

Single piping without outer insulator

300 1

1.2

1.4

1.6

Inner tube diameter or outer tube diameter (m)

Inner tube diameter or outer tube diameter (m)

(A)

(B)

1.8

Figure 5.20 Comparative study for different types of heat transport piping design. (A) Heat loss comparison. (B) Amount of material comparison.

The second is an inner insulated single piping composed of outer insulator, outer tube, liner, and inner insulator. The structural materials used in the single piping are the same as those in the coaxial hot gas duct. Results of comparative study for different type of designs are shown in Fig. 5.20. The design specification of each case is set to achieve temperature reduction of high temperature helium gas within 10 C. The liner diameter is fixed to 0.9 m to maintain pressure loss in the secondary helium cooling system below 0.2 MPa. The length of piping is set to 200 m to ensure the safe distance between reactor building and hydrogen production plant against hazardous chemical leakage

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High Temperature Gas-cooled Reactors

1.2 Cost increase by heat loss

Normalized cost

1

Construction cost

0.8 0.6 0.4 0.2 0 Coaxial hot gas duct

Insulated single piping with outer insulator

Insulated single piping without outer insulator

Figure 5.21 Comparative study for different types of heat transport piping design.

accident. Diameters of the inner tube or outer tube are determined to keep the velocity of helium gas below design target of 30 m/s. The thickness of outer insulator for the insulated single piping is set to maintain the outer tube temperature below 550 C, which is the allowable temperature for SUS316. As shown in Fig. 5.20, the coaxial hot gas duct has advantage in heat loss reduction; however, it requires larger amount of material as the size of piping increases. As for the insulated piping design, the impact of outer insulator on heat loss is significant. On the other hand, the design with outer insulator requires larger amount of material because larger thickness is required due to the temperature increase in outer tube. Fig. 5.21 depicts the overall cost including construction cost and cost increase by heat loss. The cost is normalized by the total cost of coaxial hot gas duct. As can be seen, the insulated single piping with outer insulator exhibits the minimum cost. Hence, the GTHTR300C employs the insulated single piping with outer insulator for the heat transport piping design. As for the hydrogen production plant, optimized flowsheet for the IS process is studied. The details of IS process are described in Chapter 5.6. The nuclear heat transported by the secondary helium gas cooling system is provided to the process of concentration and decomposition of sulfuric acid and hydrogen iodide. Design improvements such as introduction of membrane in hydrogen iodide decomposer and recovery of Bunsen reaction heat are made to minimize heat input from to the plant. As a result, thermal efficiency of 50.2% is achieved for the GTHTR300C [12].

5.2.3 Seawater desalination All the thermal power plants including nuclear power plants reject “waste heat” in the power generation cycle. Existing commercial nuclear plants use Rankine power conversion cycle and the waste heat becomes two third of reactor thermal power. The waste heat is rejected in the form of condensation at about 35 C, and the

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temperature is too low to be used for desalination. In order to produce portable water aside from power generation, the plant is required to choose either increasing the turbine back pressure or extract steam from a middle stage of the turbine. However, both selections result in reduction of power output in the course of nature. The Brayton power conversion system employed in the GTHTR300 does not have phase change in the cycle. As shown in Fig. 5.22, the waste heat of 316 MW in the cycle is rejected to cooling water at relatively high temperature and the water in the closed intermediate loop is elevated to 140 C. The water transports the heat to brine heater in the desalination plant and heats up the brine to about 110 C. The maximum top brine temperature is set to the temperature to avoid scale deposition. The maximum amount of the recoverable waste heat (RWH) is 248 MW. The remaining heat that cannot be recovered is rejected at the seawater cooler installed at the downstream of the brine heater. In case of outage in the desalination plant, isolation valves provided in the intermediate loop are shut to isolate the plant and the bypass valve is opened to introduce all cooling water to the seawater cooler, which is designed to have a capacity with 100% of waste heat in the nuclear plant. Fig. 5.23 shows the optimized configuration of incrementally loaded MSF concept, which consists of a brine heater, a recirculation heat recovery section, and a heat rejection section. Each stage in heat recovery and rejection sections is comprised of a condenser and an evaporative chamber. The seawater is introduced to

Figure 5.22 Intermediate water loop coupling of MSF to nuclear plant [13].

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High Temperature Gas-cooled Reactors

Figure 5.23 Concept of optimized incrementally loaded MSF process [13].

the condenser in heat rejection part and elevates the temperature by condensing the steam at the outside of the tubes. Most of the seawater is discharged with its heat; however, a part of the seawater from the heat rejection section is used as makeup for the brine. The brine exhausted from the heat rejection section is guided to the condenser in hear recovery section and increases its temperature by recovering sensible heat of steam. The nuclear heat is added at the brine heater to the desalination plant and the saturated brine liquid enters the first chamber of the heat recovery section. As it goes through the chambers, flash evaporation takes place and the salinity of the brine increases. The vapor condenses at the condenser and the distillate is collected and becomes portable water. Salinity of the portable water in the MSF process is typically smaller than that in the membrane distillation process, for example, the reverse osmosis method, by two orders of magnitude. A heat and mass balance calculation was performed for the optimized MSF process. The result showed that the amount of water production is increased by 45% over the traditional MSF process for the given amount of RWH. It is notable that the overall heat utilization rate of 88% is achieved for the reactor thermal power of the GTHTR300 by the cogeneration of electricity and water [13].

5.2.4 HTGR renewable hybrid system It is inevitable to massively reduce CO2 emission to avoid impact of climate changes and global temperature rise. A large-scale penetration of renewable energy power generation is one of the potential solutions expected in the coming decades worldwide because of the abundance and recent continual decline in cost of solar photovoltaics and wind power generators. However, the outputs of these intermittent power generations vary seasonally, daily, and hourly. The mismatches in production and consumption in electric grid induce fluctuations in its frequency and voltage, which may result in massive black out.

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Fig. 5.24 shows a concept of the tightly coupled HTGR renewable hybrid energy system [14]. The system consists of the HTGR cogeneration plant and a renewable power plant based on solar photovoltaic, wind, or both energy sources. The system integrates the HTGR and renewable systems electrically behind an electric grid at one connection point. The HTGR plant compensates the variations initiated in renewable power plants and resultant power to the electric grid is maintained constant. Fig. 5.25 shows the schematic of strategies for operation of the HTGR renewable hybrid energy system. The strategy for the economic dispatch control is to assign

Figure 5.24 HTGR renewable hybrid energy system [14].

Figure 5.25 Load-following operational strategy for HTGR cogeneration plant [14].

140 120 100 80 60 40 20 0 100

Normalized value to rated condition (%)

High Temperature Gas-cooled Reactors

Electric power Reactor power Heat supply rate to H2 plant

Power generation efficiency (%)

Power generation efficiency (%)

Normalized value to rated condition (%)

340

80 60 40 20 0

0

1 2 3 Elapsed time (h) (A)

4

140 120 100 80 60 40 20 0 100

Electric power Reactor power

80 60 40 20 0 0

1

2 3 Elapsed time (h) (B)

4

Figure 5.26 Dynamic behavior of HTGR cogeneration plant for economic dispatch control and secondary frequency control. (A) Economic dispatch control. (B) Secondary frequency control.

the heat generated in the reactor to gas turbine and industrial heat applications corresponding to the request of power generation rate from load dispatch centers. The strategy for the secondary and tertiary load frequency controls is to use the large thermal capacitance of the HTGR core as a heat storage. For an example, the GTHTR300C reactor core including fuel elements, fuel blocks, and central reflectors has thermal capacity of 370 MJ/K. Ideally, 6 MW of heat can be used to compensate the deficit of heat balance in the primary system by transferring reactor heat from and to reactor coolant. The primary frequency control can be performed by a combination of inertia of the gas turbine and a turbine bypass control, which is served as governor for gas turbine in the HTGR closed-cycle power generation system. Fig. 5.26 provides simulation results for economic dispatch control and secondary frequency control in GTHTR300C. The simulation results demonstrate that The HTGR cogeneration plant can follow the daily and hourly load variation of the renewable energy by assigning reactor heat to gas turbine and industrial plant appropriately with a combination of fundamental controls. The results also show that the HTGR can compensate the minute-scale fluctuation of the renewable energy power generation by using the large thermal capacity of the reactor core. The reactor power and high electricity generation efficiency remain constant at all times during the operation [14].

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System design for steelmaking

5.3.1 Introduction By the year 2050, CO2 emissions should be reduced from 4.8 to 1.4 Gt to limit the emissions to the half of the year 2006 [15]. The requirement is so serious that many methods should be developed. Hydrogen steelmaking can be one of such new CO2 reduction methods. Hydrogen can be used for ore reduction and only H2O is made. Hydrogen production methods from noncarbon-containing material are desired to prevent CO2 emission in hydrogen production. Thermochemical water splitting IS process is a method of hydrogen production from combination of the several reactions, which proceed thermally. SO2 1 I2 1 2H2 O ! H2 SO4 1 2HI;

(5.1)

H2 SO4 ! H2 O 1 SO2 1 0:5O2 ;

(5.2)

2HI ! H2 1 I2 :

(5.3)

The raw material of hydrogen is H2O, non carbon containing material. Heat demand of the method can be supplied from, for example, HTGR featuring supply of high temperature helium without CO2 emissions. The integration of HTGR-IS process and hydrogen steelmaking process can produce iron with theoretically no CO2 emissions. Hydrogen steelmaking also needs O2 for fuel combustion and decarbonation in electric arc furnace (EAF), etc. IS process has an advantage that O2 is produced with H2. The concepts of the NHS method applying HTGR and IS process were introduced. Heat demand and CO2 emissions of the steelmaking systems were compared with the blast furnace steelmaking (BFS) system to show the effectiveness of the concept. In addition, preliminary cost analysis of the NHS system was carried out to show the goal of nuclear hydrogen production cost.

5.3.2 Flow diagram of steelmaking systems The NHS system was analyzed through flow sheet calculation, and material balance and heat demand were compared with a BFS system. The analysis includes production and transportation of resources. Fig. 5.27 illustrates schematic diagrams of steelmaking systems of NHS, BFS, and Midrex steelmaking (MS). MS is a direct reduction steelmaking using reformed natural gas as reducing agent. In this study, the word “system” denotes the total steelmaking procedure being comprised of production of resources, transportation of resources, power generation for steelmaking, and steelmaking process itself. The detail procedure of the BF process is explained here. Coal and iron ore are transported from the mining site to the BF process. Coal is fed to the coke oven

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High Temperature Gas-cooled Reactors

Figure 5.27 Schematic diagram of steelmaking systems for heat and mass balance analysis.

(COv). Coking is made by the combustion heat of coke oven gas (COG) and blast furnace gas (BFG). Coke is produced and COG is obtained as by-product. In the sintering, ore grain is mixed with lime as flux and burned to produce sinter. Smaller ore grain is mixed with lime as flux and calcined to produce pellet. BF is used to produce pig iron. Ore, sinter, and pellet are fed to the top of the furnace with coke. Blast, which is mainly heated air, is injected to the furnace with pulverized coal and reduction reaction of iron oxides in the iron sources happens. Liquid pig iron is produced and separated from slag in the BF. BFG is blown off from the top of the furnace. The pig iron is introduced to the converter with scrap, ore, iron scale, and slag-forming fluxes. Oxygen is injected and carbon element contained in the pig iron is oxidized and removed as LinzDonawitz converter gas (LDG). Steel is tapped and the remaining material are separated to iron oxide and slag after removal and cooling. Secondary refinement of the steel is done in RuhrstahlHeraeus (RH) procedure to remove carbon and oxygen elements. High quality steel

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(HQS) of less carbon and oxygen elements is produced as the product of the process. Some part of COG, BFG, and LDG by-products gases is consumed as fuel within the process. The remaining gases are used to generate electricity by a steam gas turbine. The electricity is consumed within the process and sold to outside when excess electricity is generated. Detail of the HR process is explained. In the NHS system, uranium fuel and ore are transported to the HR process. The uranium fuel is consumed in HTGR to generate heat and electricity for the IS process and the EAF. H2 is produced in the IS process using heat and electricity from the HTGR. The H2 is fed to the shaft furnace (SF) and reduction of iron ore to direct reduced iron (DRI) takes place. Lump ore is used as iron source. The DRI, scrap, and slag forming flux are fed to the EAF. These materials are melted by electric heating. Decarburization occurs simultaneously and minute carbon element is removed as CO2 by O2 injection. The electricity and O2 supplied to the EAF are produced in the HTGR power generator and the IS process, respectively. Steel is produced and iron oxide and slag are separated as the converter in the BF process. Carbon and oxygen element impurities in the steel are removed in the RH procedure and HQS is made as the product of the process. No combustible by-product gas is produced in the process. The MS process is similar to NHS process. Natural gas and ore are transported to the MS process. Natural gas is reformed and fed to the SF. Pellet is produced from ore and flux. The pellet is reduced by the reformed gas to produce DRI. The DRI is fed to EAFs to produce steel, iron oxide, and slag [16]. The steel is refined in an RH procedure to produce HQS. Heat and mass balance of the total BF and NHS steelmaking systems were made by normalizing the production of 1 t of HQS [17]. Data sources of mass balance, temperature, and pressure of the parts are summarized in the reference [17]. Nuclear heat and electricity consumption to work the IS process to produce H2 were calculated from a material and energy analysis of the process in JAEA. A HTGR for electricityhydrogen cogeneration was applied. Nuclear heat, Q of 525.7 kJt/mol-H2 and electric energy, W of 68.6 kJe/mol-H2 including 29.5 kJe/molH2 for auxiliary such as a helium circulator for heat transfer from the HTGR were used to produce 1 mol H2. Hydrogen production thermal efficiency, η using Eq. (5.4) was 42.6%. η5

ΔHH2 ;HHV   3 100: Q 1 W=ηe

(5.4)

Higher heating value (HHV), ΔHH2 ;HHV , hydrogen combustion heat to liquid water was 285.8 kJt/mol-H2. Electricity generation thermal efficiency of the HTGR power generation, ηe was 47% assuming GTHTR300C [10]. Electricity generation by combustion of by-product gases, COG, BFG, and LDG was enough for the BF process electricity demand. Electricity generation thermal efficiency, the ratio of electricity to combustion heat of the by-product gases was set at 55.8% assuming a combined gas turbine [18]. HTGR power generation was required in the HR process because no by-product gas was generated.

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The range of heat and mass balance analysis contains the power generator, the IS process, and by-product electricity generation are in the process; the HTGR is not included. Heat input is the sum of the heat from the HTGR and combustion heat of the coal. Heat input for the production and transportation of ore, coal, and uranium was considered. Other materials such as lime were not taken into consideration because relative amount of them was less than the materials considered. And heat input in electricity generation procedure itself was also considered. Data of the heat input were summarized in the reference [17]. Table 5.8 shows the mass balance of the steelmaking processes. Although iron source inputs of the BFS system were at 1802 kg/t-HQS, iron source input of the NHS system was 1696 kg/t-HQS. Some iron element was moved to the slag in the BF process. More iron source was needed to compensate the loss. More scrap input, which contained less oxygen element, was another reason for the less iron source in the NHS system. For reducing agent/fuel, 755 kg/t-HQS coal was required for the BFS system. On the other hand, only uranium of 4.8 3 1022 kg/t-HQS was needed for the NHS system due to the high energy density of the nuclear fuel. Smaller burden of transportation is expected in the NHS system. Flux input of 44 kg/t-HQS in the NHS system was much less than that of 212 kg/t-HQS in the BFS system. Output slag was also smaller in the NHS system as the difference of the flux input. The difference was mainly from the difference of the higher ratio of sinter and pellet input to BF than lump ore because sintering and palletization required flux. Table 5.9 shows the heat input to the steelmaking systems. As in Table 5.9A, fuel of 22.1 GJt/t-HQS was input and excess electricity of 4.7 GJt/t-HQS was obtained from the BF process. Power generation from by-product was more than the requirement of the BF process. HTGR nuclear heat of 17.1 GJt/t-HQS and electricity of 4.0 GJt/t-HQS were input to the NHS process. Table 5.9C shows that net heat input (difference of heat input and output) of the BF process of 13.8 GJt/t-HQS was the smaller due to the high efficiency of coal combustion. Lower thermal

Table 5.8 Mass balance of the steelmaking systems. Material

Input Iron source (kg)

BFS system NHS system

Other (kg)

Ore

Scrap

Scale

Coal

Uranium

Flux

1688 1589

71 107

43 

755 

 4.8 3 1022

212 44

Intermediate

Material BFS system NHS system

Reducing agent/fuel (kg)

Output

Reducing agent/ fuel (Nm3)

Product (kg)

Hydrogen  727

HQS 1000 1000

By-product (kg) Iron oxide 50 39

Slag 409 188

Exhaust (Nm3) CO2 1047 7

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Table 5.9 Heat input to steelmaking systems including production and transportation of material and power generation. BFS system

NHS system

(A) Electricity and heat input to steelmaking process, unit: GJt/t-HQS (heat), GJe/t-HQS (electricity) Steelmaking process Fuel HTGR Power generation/excess electricity

22.1  2 4.7

 17.1 4.0

(B) Heat input to steelmaking process and other procedures, unit: GJt/t-HQS Steelmaking process Fuel HTGR Power generation/excess electricity (converted to heat)a Material production Material transportation Power generation

22.1  2 8.4 0.8 3.1 0.0

 17.1 8.5 0.5 2.3 0.1

13.8 3.9

25.5 2.9

17.6

28.4

(C) Net heat input, unit: GJt/t-HQS Net heat input to Steelmaking Net heat input to material production, material transportation and power generation Total net heat input a

Converted to heat with power generation efficiency of 55.8% considering combined gas turbine [18] for the BFS system, and that of 47% considering the HTGR power generation [10] for the NHS system.

efficiency of IS process hydrogen production compared with coal combustion was a major reason of the large net heat input of 25.5 GJt/t-HQS. However, there is a limitation of improvement of the hydrogen production efficiency. Thermal efficiency of IS process of 42.6% was already high estimation compared with other estimations [19]. Reduction of hydrogen consumption for combustion within the SF is desired. Heating H2 before injection to the SF by combustion of fossil fuel is one idea. Of course, the balance of net heat input and CO2 emissions should be considered. Waste heat from the EAF (1.3 GJt/t-HQS) was another reason of the large net heat input. Improvement of the EAF is also necessary. Table 5.9B shows the heat demand input for material production, material transportation, and power generation. Total amount of them was 3.9 GJt/t-HQS, and 2.9 GJt/t-HQS in the BFS system, and NHS system, respectively. Much less amount of uranium than coal had an effect on the less demand of the NHS system. Total net heat demand was shown in Table 5.9C. Heat of 17.6 GJt/t-HQS and 28.4 GJt/tHQS was resulted for the BFS system and NHS system, respectively. The advantage of less heat demand for transportation, etc. was not so large to change the order of total net heat input.

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5.3.3 CO2 emission CO2 emission from the BF and NHS systems was estimated and compared to evaluate effectiveness of NHS on CO2 emission reduction. CO2 emission from steelmaking processes was evaluated from material balance made in the previous section. CO2 from gas turbine power generation by combustion of by-product gases was taken into consideration. CO2 emissions for the production and transportation were considered for ore, coal, and uranium. In addition, CO2 emission from electricity generation itself was also considered. Data of the CO2 emissions are summarized in the reference [17]. Table 5.10 shows the CO2 emissions. Though CO2 from the BF process was 1047 Nm3/t-HQS, the one from the HR process was only 7 Nm3/t-HQS. Carbon element was fed to the process only from electrode in the EAF. Total CO2 emissions from material production, material transportation, and power generation were 137 Nm3/t-HQS and 103 Nm3/t-HQS in the BFS system and NHS system, respectively. Just as the discussion of the heat input, small amount of uranium fuel had effect on the less CO2 emissions. Total CO2 emissions from the NHS system (110 Nm3/t-HQS) was 9% of that from the BFS system (1184 Nm3/t-HQS). The result shows the NHS applying HTGR and IS process is effective for reduction of CO2 emissions.

5.3.4 Steelmaking cost Cost of steelmaking by NHS was evaluated by comparing with BFS and MS. Estimation of the HTGR-IS process hydrogen production cost at which the NHS system was competitive to the BFS and MS systems was the purpose of the evaluation. Steelmaking cost was considered as sum of the cost of materials, plant, operation, and carbon capture and storage (CCS). Material cost was also divided into the following categories: iron sources (ore, pellet, sinter, and scrap), reduction agents (coke, natural gas, and hydrogen), other materials (lime, carbon electrodes, etc.). Material cost included mining, transportation, and production (coking, sintering, pelletization, and HTGR-IS process) cost of them. Plant cost meant the cost of plant facilities: a natural gas reformer, BF, SF, Converter, EAF, RH, ancillary facilities, and by-product power generation and the generated power from it. Operation cost meant the one to operate the facilities, such as electricity and labor. CCS cost meant the cost to recover CO2 Table 5.10 CO2 emissions from steelmaking systems including production and transportation of material and power generation [17] (unit: Nm3/t-HQS).

Steelmaking process Material production Material transportation Power generation Total Total system

BFS system

NHS system

1047 25 112 0 137 1184

7 16 85 2 103 110

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from the process and the one to transport and store the CO2. Main assumptions in the cost analysis were as follows: The data used to calculate the steelmaking cost were mainly taken from a report [7]. Cost data are summarized in Ref. [17]. G

G

G

G

G

G

G

G

The range of evaluation was from mining of materials to HQS after secondary refinement. Steelmaking processes were located in midland of the United States. Scales of both the steelmaking plants were fixed to one million t-HQS/y. This scale was small compared with actual steelmaking plants. Plant cost per unit HQS production tends to be less expensive in large scale. Therefore plant cost was estimated to be more expansive than actual plants. Process plant components, heat, and mass balance were different from the processes considered for material, heat balance, and CO2 emission. Heat and energy requirement of the IS process was also different. For iron ore, coal, and natural gas, low cost case in year 2000 and high cost case in year 2008 were analyzed to consider the effect of the costs of them. Cost of other materials such as uranium, electricity, and labor was fixed. Material cost of hydrogen from HTGR-IS process was taken from the literature [20]. This cost datum was rather inexpensive compared with other HTGR-IS cost estimations. Because of the uncertainty of the cost, effect of hydrogen production cost on the steelmaking cost of the NHS system was investigated. Cost of the NHS system was calculated using cost data of the MS system [21] as a basis. Electricity used in the BFS system and MS system was generated from coal power plant. On the other hand, electricity for NHS system was generated from HTGR, which did not emit CO2. Year of cost data was unified to year 2000 by using consumer price index in the United States [22].

Steelmaking cost (US$/t-HQS)

Fig. 5.28 shows the steelmaking cost. Cost of plant, iron source, reduction agent, and CCS had influence on the difference of the steelmaking cost: cost of operation and other materials had less effect. The difference of the iron source cost in the

800

640

700

73 32 46 49 14

73 34 100 50 18

700 600 500 400

625

50 57 151

100

426

50 98

48 37

282

282

426

0

Price is modified to that in year 2000.

Figure 5.28 Steelmaking cost [17].

618

678

CCS 0 151

211

48 37

300 200

727

151

0 151

211

47 37

48 37

232

232

Material (reducti on agent) Material (iron source) Material (others) Operation

Plant

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BFS system and the other systems was mainly made from the difference of iron source. While lump ore, sinter, and pellet were used in the BFS system, only pellet was used in the MS and NHS systems. Pellet cost itself was different between the BFS system and the other systems in the literature data. The difference of the cost of reduction agents was made from the difference of the agents: coke for the BFS system, natural gas for the MS system, and hydrogen for the NHS system. In the year 2000, the low material cost case, the order of steelmaking cost is the BFS system (640 US$/t-HQS), the MS system (625 US$/t-HQS), and the NHS system (618 US$/t-HQS). The main factors of the cost in these systems were different. Plant cost was highest in the BFS system. Though plant cost was less expensive in the MS system, iron source was of high cost. The plant cost was lowest in the NHS system. Also, no CCS cost was used. On the other hand, hydrogen cost as a reduction agent was highest. Hydrogen production cost and CCS cost were important factors for competitiveness of the NHS system. In the year 2008, the high material cost case, the order of BFS system and MS system was exchanged. The order was the MS system (727 US$/t-HQS), BFS system (700 US$/t-HQS), and the NHS system (678 US$/t-HQS). Only the material cost was influenced by the change. Increase of the cost was great in the MS system, where ratios of iron source and reducing agent cost were larger. Cost increase was less in the NHS system because hydrogen cost was not changed. For the NHS system, competitiveness would be expected if material cost of hydrogen was not changed in 2008. Here, it is stressed that material cost of hydrogen was the same in both the cases. Actually, cost of uranium became about 4 times in year 2008 compared with that in year 2000 [23] and the ratio of fuel cost in the nuclear electricity cost cannot be negligible. To evaluate the future competitiveness of the NHS system, breakdown of the HTGR-IS process hydrogen cost is necessary. Fig. 5.29 shows the effect of hydrogen production cost on the NHS cost in the low material cost case in 2000 and high material cost case in 2008, (B) 800

Steelmaking cost (US$/t-HQS)

Steelmaking cost (US$/t-HQS)

(A)

700

600

500

400

0

1

2

3

4

Hydrogen production cost (US$/kg-H2)

800

700

BFS system MS system NHS system

600

500

400 0

1

2

3

4

Hydrogen production cost (US$/kg-H2)

Price is modified to that in year 2000.

Figure 5.29 Effect of hydrogen production cost on steelmaking cost. (A) High material cost case in 2008. (B) Low material cost case in 2000 [17].

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respectively, to evaluate the hydrogen production cost on the steelmaking cost. The cost was compared with the steelmaking cost in the BFS and MS systems. In the low material cost case (Fig. 5.29A), the NHS steelmaking cost at hydrogen production cost of 3.3 US$/kg-H2 was the same as the BFS steelmaking cost. When hydrogen production cost was 3.0 US$/kg-H2, the NHS steelmaking cost was the same as the MS steelmaking cost. In the high material cost case (Fig. 5.29B), the hydrogen production costs of 3.3 and 3.8 US$/kg-H2 were of the equivalent steelmaking costs of NHS and BFS, and NHS and MS, respectively. Higher hydrogen cost is allowed for the competitiveness of the NHS system to the MS system. On the other hand, hydrogen cost required for the competitive NHS system to the BF system was not changed. The result shows cost objective of the NHS system. However, more accurate data of the HTGR-IS process hydrogen production cost including uranium fuel cost and influence of scale effect of steelmaking plants were necessary for further consideration whether the objective of the NHS system can be attained.

5.4

Safety design for connection of heat application system and high temperature gas-cooled reactor

5.4.1 Introduction HTGR is expected to extend the use of nuclear heat to nonelectric heat applications such as hydrogen production, process heat supply, due largely to high temperature heat supply capability as well as inherently safe characteristics. Such characteristics highly depend on the use of ceramic-coated fuel particle, graphite moderator, and helium coolant. Toward the commercialization of HTGR, safety standard should be established fully taking into account the inherent characteristics of the HTGR. In addition, the standard for coupling heat application system, for example, hydrogen production plant should be developed. This chapter mainly describes activities related to establishment of safety standard for commercial HTGR in Japan.

5.4.2 Roadmap for safety standard establishment Fig. 5.30 shows a roadmap of a safety standard establishment for licensing of commercial HTGR cogeneration systems in Japan. Prior to the construction of the HTTR, the safety standards were developed based on those of LWR [24]. Although several unique features of HTGR were incorporated and excess requirements, which were specific to LWR, were eliminated, some of the safety characteristics of HTGR could not fully taken into consideration because the construction and operation was the first case in Japan. In addition, the standards did not cover the requirements for coupling heat application system to nuclear reactor because the HTTR was not designed to provide heat to the heat application system at the time. In order to develop safety standards for commercial HTGR cogeneration system, a research committee on the safety standards for HTGRs was established in 2013 under the Atomic Energy Society of Japan (AESJ) with the support of the Research Association of

350

High Temperature Gas-cooled Reactors AESJ: Atomic Energy Society of Japan

Active safety

Passive safety Inherent safety

HTTR safety standards LWR safety standards Basic research Seismic evaluation methods Safety analysis data

HTTR safety standards, safety evaluation methods

HTTR

1998

Future HTGR

HTTR safety demonstration test

Data acquisition on inherent Commercial HTGR safety features of HTGR (electricity, H2) Reactivity insertion Coolant flow reduction Loss of forced cooling (30%) Fuel performance

HTTR data acquisition

Present

HTTR HTTR-GT/H2 plant

Drafting safety requirements The research committee under AESJ (FY13-14) Requirements for commercial HTGRs Requirements for coupling H2 facility to HTGR

2020

Drafting safety requirements Drafting safety guides in AESJ (FY15 - 16) IAEA CRP (Dec. 2014 - )

2030

2040

Figure 5.30 Roadmap for establishment of safety standards for HTGR cogeneration system [24].

HTGR safety features HTTR data

Safety approach

Drafting of safety requirements (FY13 - FY14)

Safety fundamental

Safety requirements

Safety guides

Establishment of basic concept of safety guides (FY15 – FY16) Safety design Evaluation items Design basis events Acceptance criteria System design

Future activity Safety standards

Safety design process

Figure 5.31 Activities for development of safety standards in AESJ [24].

High Temperature Gas Cooled Reactor Plant (RAHP). The activities in the committee are shown in Fig. 5.31. The committee aimed to develop safety standards for plant design, which is unique to HTGR, which can be the basis for international standards. In the first two years, safety requirements were drafted considering the HTGR safety features demonstrated by the HTTR, and lessons learned from the TEPCO Fukushima Daiichi nuclear power plant accident. In addition, requirements for integration of heat application system (e.g., hydrogen production plant) to HTGR were also developed. The following two years were dedicated for developing a basic concept of safety guides. The following section provides brief summary of outputs from the committee.

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5.4.3 Safety requirements 5.4.3.1 Basic safety approach The following basic safety approaches are employed for the development of safety requirements. Practically eliminate plant conditions, which result in massive releases of radionuclides from the plant to environment. Protect integrity of radionuclide retention barrier of coated fuel particle. Maintain required safety functions relying on inherent and passive safety features.

G

G

G

5.4.3.2 Safety requirements With the aim of international standardization, IAEA Specific Safety Requirements for LWR design (SSR-2/1) [25] are used as the basis. Table 5.11 shows major differences in the safety requirements between LWR and HTGR. The details are summarized in the following sections.

Table 5.11 Comparison of safety requirements between HTGR and LWR [24]. Safety requirements

HTGR

LWR

Confinement of radionuclide Physical barriers Protection of fuel integrity Containment vessel or confinement Control reactivity G

G

G

Heat removal from core

Loss-of-offsite power

Set of barriers (functional containment) In operational states and accident conditions Confinement

At least two divers and independent means (inherent design features are regarded as one of means) Passive cooling from outside surface of reactor vessel (passive cooling) Monitor plant parameters and radiological conditions (emergency power supply is not required to mitigate the consequences of AOOs and accidents)

Multiple barriers In operational states (normal operation and AOO) Containment vessel

At least two divers and independent means

In shutdown states: residual heat removal (forced cooling) In accident condition: emergency core cooling (forced cooling) Mitigate the consequences of AOOs and accidents, monitor plant parameters and radiological conditions

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5.4.3.2.1 Confinement of radionuclides Fuel integrity is only required during normal operation and anticipated operational occurrences (AOOs) in the LWR. On the other hand, the retention capability is required to be maintained even in accident conditions for the HTGR to alleviate airtightness limits for confinements.

5.4.3.2.2 Control reactivity Autonomous reactor shutdown without exceeding fuel temperature limit is possible in HTGR even in case of loss of forced cooling accidents with failures of all reactivity control systems. Such characteristics is due to the unique elements of refractory coated fuel particles, large negative temperature coefficients, larger heat capacity, low power density in core, and allow the HTGR safety requirements to regard inherent design features as one of means for safety function of “Control reactivity.” The adequacy of the requirement has been demonstrated by safety demonstration tests using the HTTR.

5.4.3.2.3 Heat removal from core HTGR can remove core heat passively independent of reactor coolant existence and therefore core decay and residual heat cooling systems by forced convection inside the reactor are not needed under all circumstances. Because of such characteristics, the HTGR safety requirements mandate to cool the reactor passively and externally at the outside surface of reactor vessel for the safety function of “Remove core heat.”

5.4.3.2.4 Loss-of-offsite power Emergency power supply for mitigation of the consequences for AOOs and accidents is not required in HTGR according to the basic safety approach. The requirement only mandates to supply power for monitoring plant parameters and radiological conditions.

5.4.3.2.5 Coupling to heat application system Heat application system may pose the following events when coupling to HTGR: G

G

Temperature and pressure transients initiated by abnormal events in heat application systems. Abnormal transportation of hazardous chemicals to nuclear facility.

The fluctuations of process values in the nuclear facility are already considered in the HTGR safety requirement. The transfer of hazardous materials in industrial plants located in the vicinity of nuclear facility is also considered in the existing safety requirements for external events. However, the amount of chemicals such as hydrogen involved in heat application plants connected to HTGR would be large. In addition, the pipping, which integrates a nuclear facility and heat application system, may become a new pathway for the transportation of chemicals to nuclear facility. Hence, the discussion in the committee concluded that the following requirements should be stated explicitly: G

G

G

Protect reactor coolant pressure boundary and confinement against fire and explosion of combustible chemical leaked from heat application system. Mitigation of corrosive chemical intrusion into reactor coolant pressure boundary. Mitigation of toxic chemical intrusion into control room.

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Another important outcome that is derived in the committee is the certification requirement, which allows the heat application system to designate as nonnuclear facility in order to comply with a request from potential users. The requirements are as follows: G

G

Maintain normal operation of nuclear facility during every possible condition in heat application system. Mitigate tritium concentration in heat application system to allowable level.

5.4.4 Basic concept of safety guides 5.4.4.1 Evaluation items Evaluation items for the design of reactor core, reactor coolant system design, secondary coolant system design, decay heat removal system design, and confinement design to meet the safety requirements of fuel design are investigated. Table 5.12 provides the results for fuel design as an example. First, failure modes, which may lead to loss of safety function, are investigated. Second, mechanisms that cause the failure modes in each operational state are derived. Finally, figure of merit corresponding to the evaluation items is obtained.

Table 5.12 Evaluation items and figure of merits for fuel design [24]. Functional requirements

Evaluation items to meet safety requirements

Figure of merits

Structural integrity during normal operation

(a) Internal pressure increase due to generations of fission product and free oxygen (b) Kernel migration due to free oxygen generation and temperature gradient (a) Thermal degradation of SiC layer due to abnormal temperature increase (b) Internal pressure increase due to abnormal temperature increase (c) Oxidation due to oxidized gas (O2, H2O)

(a) Stress induced by internal pressure

Structural integrity during AOOs and accidents

Structural integrity during fuel handling

(a) Load and stress during fuel handling

(b) Distance of kernel

(a) Fuel temperature

(b) Stress, enthalpy

(c) Index of corrosion circumstance (e.g., partial pressure of oxygen and temperature, etc.) (a) Stress

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High Temperature Gas-cooled Reactors

5.4.4.2 Licensing basis event selection Licensing basis events (LBEs) should include multiple failure events to be consistent with the safety requirement. First, accident sequences which are a set of initiating events followed by a sequence of failures in safety function. Second, significant accident sequences are selected from the accident sequences in reference to acceptance criteria (cf. Section 5.4.4.3) used in the safety analysis. The events whose potential consequence or frequency is sufficiently small can be screened out in the process of selection. However, the continuous monitoring for the consequence and frequency should be conducted to reflect the latest findings. The LBE selection procedure based on the approach is as follows:

5.4.4.2.1 Identification of abnormal events and postulated initiating events Abnormal events and postulated initiating events (PIEs), which lead to abnormal release of radioactive materials from the plant to environment, are identified. A systematic approach such as Master Logic Diagram (MLD) and Failure Mode Effect Analysis (FMEA) should be used for the identification.

5.4.4.2.2 Definition of safety function and mitigation system Safety functions and corresponding mitigation systems against abnormal release of radioactive materials in case of abnormal events are defined and classified into the following categories: G

G

G

G

Control reactivity Heat removal from core Control chemical attack Confinement of radionuclides

5.4.4.2.3 Grouping of abnormal events Abnormal events are binned to “PIE group” in terms of similarity of mitigation systems expected in the accident progressions. The binning helps in identification of accident sequences.

5.4.4.2.4 Licensing basis event selection for single failure events PIEs are classified into AOOs and accidents depending on the event occurrence frequencies. Representative events can be selected from each PIE group in case that the events lead to the most severe consequence in the evaluation items corresponding to acceptance criteria.

5.4.4.2.5 Identification of accident sequence Accident sequences for selected PIE group are identified by event tree analysis. Mitigation systems defined in step b should be used as headings in accident sequences.

5.4.4.2.6 Grouping of accident sequence Accident sequences are binned in terms of similarity of PIEs and accident progressions. The binning facilitates identification of significant accident sequences.

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5.4.4.2.7 Identification of significant accident sequence Significant accident sequences, which result in the most severe consequence in the evaluation items corresponding to acceptance criteria, are identified. Accident sequences whose frequencies or consequences are sufficiently small can be eliminated.

5.4.4.2.8 Licensing basis event selection Select single failure events and significant accident sequences identified earlier as LBEs.

5.4.4.3 Acceptance criteria Figure of merits identified in the identification process of “evaluation items” can be used as the parameter for the acceptance criteria. The thresholds for the acceptance criteria are determined based on design codes, test data, etc., considering the conservativeness. The following are the examples of the acceptance criteria based on the knowledge of HTTR design:

5.4.4.3.1 Anticipated operational occurrences G

G

Peak fuel temperature shall not exceed 1600 C. Integrity of reactor coolant pressure boundary shall be assured. Pressure on reactor coolant pressure boundary shall not exceed 1.1 times the maximum working pressure. Temperature of reactor coolant pressure boundary shall not exceed design limit. G

G

5.4.4.3.2 Accidents G

Coated fuel particle shall not significantly fail, and reactor core shall not be seriously damaged and its coolable geometry shall be maintained. Coated fuel particle shall not fail to the level that public dose become significant. Fuel region in fuel element shall remain in graphite block. Support column shall maintain its strength to support reactor core. Integrity of reactor coolant pressure boundary shall be assured. Pressure on reactor coolant pressure boundary shall not exceed 1.2 times the maximum working pressure except the boundary between primary and secondary coolant systems. Boundary between primary and secondary coolant systems shall not fail. Temperature of reactor coolant pressure boundary shall not exceed design limit. Integrity of confinement shall be assured. Confinement boundary shall not fail by pressure loading. Temperature of confinement boundary shall not exceed design limit. Combustible gas concentration in confinement shall be out of explosive range. Risk of public dose shall not be significant. G

G

G

G

G

G

G

G

G

G

G

G

5.4.5 HTTR cogeneration demonstration In order to demonstrate inherent safety features of HTGR as well as provide valuable data on coupled neutronic/thermal fluid transients for code validation, the safety demonstration tests have been performed. Details of the HTTR design and

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High Temperature Gas-cooled Reactors

the safety demonstration test can be found in Chapter 2, Design of High Temperature Engineering Test Reactor and Chapter 4, Operation of High Temperature Engineering Test Reactor. The HTTR is also suitable for a test bed for nuclear cogeneration demonstration. The demonstration is necessary to establish the safety standards for coupling heat application system to the nuclear reactor through the receipt of reactor installation license from Nuclear Regulatory Authority. The demonstration can also contribute to establish design considerations of GTHTR300C, a commercial HTGR cogeneration plant, providing compliance with the safety requirements. One of the key items to be demonstrated for an HTGR helium gas turbines is suppression of turbine over speed against loss of generator load. HTGR closedcycle gas turbines are not able to vary heat supply rate immediately because time constant of the reactor is large. Isolation valves, which are typically used to isolate the reactor from power conversion system, cannot be employed in the direct cycle HTGR plant because of the high temperature. The approach for the turbine speed control in HTGR is to adjust cycle pressure ratio by means of controlling bypass flow to transfer a part of helium from the high-pressure side of the cycle to lowpressure side. The control not only reduces shaft power in turbine but also increases compression work in compressor and results in the reduction of turbine speed. The licensing is also expected to be utilized for establishment of design considerations for coupling chemical plants, that is, hydrogen production plant. Examples of the design considerations for the assurance of reactor safety are assurance of a safe distance and installment of isolation valves between the nuclear facility and chemical plant, etc. A control system for turbine inlet temperature that can be elevated in case of loss of thermal load in a chemical plant is an example to designate chemical plant as nonnuclear facility. Fig. 5.32 depicts the system layout of the HTTR gas turbine cogeneration plant (HTTR-GT/H2 plant) designed by JAEA [26]. Fig. 5.33 shows the bird’s eye view of the plant. The HTTR-GT/H2 plant incorporates those design considerations. In order to reduce the project cost due to the extensive modification in existing nuclear facility, the helium gas turbine and second intermediate heat exchanger (second IHX), which provides heat to hydrogen production plant, are installed in secondary helium cooling system. The second IHX and helium gas turbine are installed in series to simulate cascade heat application of the GTHTR300C [27]. Table 5.13 shows the major plant parameters. The HTTR operates in parallel operation mode and one-third of the reactor heat is transferred to the secondary loop at the IHX. In cogeneration operation, the helium elevated the temperature to 900 C at the IHX and provided to the second IHX at 887 C through the inner side of coaxial hot gas duct and inner-insulted piping. The helium exhausted from the second IHX is then guided to the helium gas turbine at 568 C. On the contrary, the helium is directly introduced to the gas turbine at 568 C in sole power generation operation. The turbine exhaust heat is recovered in the recuperator. Then the helium is delivered to a precooler, enters a compressor, reheated at the recuperator, and returns to the IHX. A cooler is installed upstream of IHX to maintain the temperature of helium flowing back to the IHX. The return temperatures are 374 C and

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H2 plant

Isolation valve

2nd IHX Mixing volume Compressor

Vessel cooling system

Auxiliary water air cooler

Turbine

Generator/motor

Auxiliary heat exchanger

Precooler

Recuperator

Cooler Air cooler Reactor

Intermediate heat exchanger

Primary pressurized water cooler

Pressurized water air cooler

Cooling tower

Figure 5.32 HTTR-GT/H2 plant system layout [26]. Reactor auxiliary system Gas turbine

Integrated heat exchanger

2nd IHX

Reactor Building

HI distillation

H2SO4 decom poser

HI decomposer EED

Isolation valves

Bunsen reactor

Reactor Cooling water system

PPWC CV IHX

Figure 5.33 HTTR-GT/H2 plant bird’s eye view [27].

150 C in sole power generation and hydrogen cogeneration operations, respectively. The recuperator and precooler are installed in a single vessel to simulate the configuration of heat exchanger unit in GTHTR300C. Table 5.14 provides a set of demonstration items, test items, and design target for the HTTR demonstration program. Based on the experience in a design of test facility for helium gas turbine, cycle pressure ratio above 1.3 is required to have transients in loss of generator load test. The turbine design incorporates two improvements to solve issues raised by the employment of radial type turbines, which can be only procured

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Table 5.13 Major specifications of HTTR-GT/H2 plant. Reactor operation mode

Reactor power (MWt) IHX heat rate (MWt) PPWC heat rate (MWt) Reactor outlet temperature ( C) Turbine inlet temperature ( C) Turbine flow rate (kg/s) Compressor pressure ratio Second IHX heat rate (MWt) Power generation (MWe) Hydrogen production (Nm3/h)

Sole power generation

Cogeneration

Rated operation

High temperature operation

30 10 20 850 568 9.7 1.4 0 0.7 0

30 10 20 950 568 9.0 1.4 0.7 0.2 30

Table 5.14 Demonstration items, test items, and design target for HTTR-GT/H2 plant. Demonstration items

Test items

Design targets

Start-up and shutdown operation procedure for direct cycle HTGR helium gas turbine system Control method for electricity load following operation

Start-up and shutdown operation simulating direct cycle in HTTR-GT/H2 plant Load-following operation test The procedure is a part of the start-up and shut down operation Hydrogen production plant load following operation test

Determine start-up and shut down sequence

Control method for hydrogen production plant load following operation

G

Control method against loss of generator load

Loss of generator load test at sole power generation mode in HTGR-GT/H2 plant

Control method against upset of hydrogen production plant

Simulation test for upset of hydrogen production plant at cogeneration mode at cogeneration mode in HTTR-GT/H2 plant

Simulating direct cycle HTGR helium gas turbine system Determine heat and mass balance allowing sole power generation during loss of thermal load in hydrogen production plant Determine system heat and mass balance allowing observation of plant dynamic behavior during loss of generator load Determine system heat and mass balance allowing observation of plant dynamic behavior during simulation test for upset of hydrogen production plant

domestically. First, a turbine, which has expansion ratio of 1.14 at optimum operational range, is employed and connected in series rather than using a turbine with expansion ratio of 1.3. Second, two turbines are devised at either end of a generator shaft in order to avoid the shaft vibration initiated by torsional torque generation. The design

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improvements not only increase the cycle pressure ratio but also increase the heat supply rate to hydrogen production plant from 0.07 to 0.7 MW. The increase is mainly due to the reduction of waste heat in the precooler and cooler because of the decrease in exhaust temperature at turbine. The increase in heat supply rate allows HTTR-GT/H2 plant to observe temperature transients during the simulation test for upset of hydrogen production plant at cogeneration mode in HTTR-GT/H2 plant. As a result of heat and mass balance evaluation and manufacturability assessment by an equipment supplier for the HTTR-GT/H2 plant design, it was confirmed that demonstration items can be validated in the HTTR-GT/H2 plant with key achievements: hydrogen cogeneration with power output 0.7 MWe at cycle pressure ratio above 1.3 and hydrogen production rate 30 Nm3/h.

5.5

Gas turbine technology for power generation

5.5.1 Introduction JAEA conducted R&D on the power conversion system of the GTHTR300 plant shown by Fig. 5.5, in parallel with plant design work [2]. The design of the power conversion system is based on a regenerative, nonintercooled, closed Brayton cycle with helium gas as the working fluid. A single-shaft, axial-flow turbo-compressor and a directly coupled electric generator run on journal and trust bearings as shown in Fig. 5.9. Magnetic bearings will be used as journal and trust bearings. The major R&D issues for the power conversion components are a helium gas compressor and a high load capacity magnetic bearing. The test plans were set up to address these issues, aiming at verifying the design of the GTHTR300 power conversion system and establishing key technologies of a closed-cycle helium gas turbine system [28]. The compressor aerodynamic performance test is aiming at verifying the aerodynamic performance and design method of the helium compressor. An earlier experience in a closed-cycle helium gas turbine was a 24,000-h operation of the EVO cogeneration facility [29] in Germany for 19741988, but it failed to achieve the design power output of 50 MWe. In 1981 a large-scale turbo-machine was built and put on tests in the German HHV facility [29]. After a 325-h operation of the HHV turbo-machine at 850 C, the facility was shut down due to the termination of the high temperature helium turbine (HHT) project in Germany by the German government. JAEA fabricated a 1/3-scale, four-stage compressor test model and a helium gas loop. The model was fabricated to simulate the repeating stage flow, and at the same time have satisfactorily high machining precision, Reynolds number, and measurement accuracy. The helium gas operating pressure is varied to investigate the effects of the Reynolds number on the efficiency and surge margin. Two sets of blades were fabricated to evaluate the effects of the end-wall over-camber angle. Test results will provide the basis for further improvement in the GTHTR300 compressor design [7,30]. The magnetic bearing development test is aiming at developing the technology of the magnetic bearing supported rotor system. For GTHTR300, the rotor mass is 52 t for the turbo-compressor and 67 t for the generator; each rotor is supported by

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High Temperature Gas-cooled Reactors

two radial bearings. Under normal operating condition, the turbo-machine rotor runs at a commercial frequency of 60 Hz (3600 rpm). Because of the low stiffness of magnetic bearings, the turbo-compressor and generator rotors are inevitably designed as flexible rotors that have bending mode critical speeds below the rated rotational speed. Presently, the heaviest rotor that is supported by magnetic bearings is the 35-t waterwheel rotor at the Yamazaki hydraulic power station of Tokyo Electric Power Company. The rotor turns rather slowly at a rated speed of 600 rpm, and is termed as a rigid rotor that does not exceed the critical speed of bending mode. On the other hand, the heaviest flexible rotor that is supported by magnetic bearings is the 9.3-t electric motor that drives a natural gas compressor rotor of NAM in the Netherlands [31]. The rotor is supported by three radial magnetic bearings and turns in the speed range from 600 to 6300 rpm traversing the critical speed of bending mode. The heavy rotors of the GTHTR300 turbo-machine require much higher load capacity magnetic bearings than now available. They also put a heavy demand on the auxiliary bearings. Magnetic bearings together with auxiliary bearings need development to scale-up for the GTHTR300 application. JAEA made a demonstration test program with 1/3-scale turbo-compressor and generator rotor models connected together by a flexible coupling. Each rotor model is supported by two radial magnetic bearings with a high load capacity that is about 1/10 of the GTHTR300 design. Prior to the demonstration test, JAEA designed a control system for magnetic bearings with MIMO controller, which linked magnetic forces of multiple magnetic bearings by feedback of multiple measurement values of vibration of a rotor, was applied for the magnetic bearing suspending the generator rotor to restrict the whirling motion [32]. The demonstration test is left as one of the future development issues.

5.5.2 Helium gas compressor JAEA’s 300 MWe class nuclear helium gas turbine design is shown in Fig. 5.34. The unit consists of a turbine and a compressor on single shaft. The turbine bell mouth intakes hot helium ducted from the reactor, whereas the compressor radial diffuser discharges high pressure helium back to the reactor. The turbine exhaust and compressor inlet scroll casings in the unit midsection connect to the closed Brayton cycle heat exchangers. Bearings are located to the rotor ends. The shaft cold end drives a 3600 rpm synchronous generator directly. The unit’s structural

Figure 5.34 300 MWe class helium gas turbine design [7].

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design follows largely the practice of industrial air gas turbines in that it employs similar rotor disks designed to comparable stresses, similar rotor orientations and bearing spans, and similar casings. The test compressor shown in Fig. 5.35 includes four stages of rotor and stator blade rows in the main bladed flow path, a row each of inlet and outlet guide vanes, and inlet and discharge scroll casings. Four stages are selected for two reasons: (1) to observe end wall boundary layer growth through multiple rotating blade rows and (2) to test blading performance in a stage sufficiently removed from machine entry and exit effects. The main bladed flow path models the forward stages of the full compressor and comes equipped in two sets of airfoils, referred to as Case 1 and Case 2, with different cambers near end walls. Table 5.15 compares principal design values of the test and full-scale compressors. By employing the conventional design method of blade, inlet blade angle mismatched with flow angle around the end walls of hub and tip as thickness of boundary layer increased in every stage. In the boundary layer, inlet flow angle increased because axial velocity decreased. It is necessary to increase the inlet blade angle at the leading edge for matching with the flow angle around the end wall of hub and tip of blade. In the aerodynamic tests of model compressor, two cases of blade were tested, one is a 1/3-scale nominal design blade of the actual compressor, which has

Figure 5.35 Helium test compressor in one-third dimensional scale [7].

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High Temperature Gas-cooled Reactors

Table 5.15 Outline design parameters for full and test compressors [7]. Full compressor

Test compressor

3.52 28.4 2.0 44.2 3600

0.883 30 1.15 12.2 10,800

20 1.70 1.50 321 72/94 78/60 1.19/1.20 1.3/1.7 B1% blade span 0.51 0.63 High reaction

4 0.57 0.50 321 71/94 26/20 1.19/1.20 1.3/1.7 B1% blade span 0.51 0.63 High reaction

Nominal design conditions Inlet pressure (MPa) Inlet temperature ( C) Pressure ratio, flange to flange Mass flow (kg/s) Rated speed (rpm)

Aerodynamic design pitch line value Number of stages Tip diameter (first rotor, m) Hub diameter (first stator, m) Tip speed (first stage, m/s) Rotor/stator blade count (first stage) Rotor/stator chord (first stage, mm) Rotor/stator solidity (first stage) Rotor/stator aspect ratio (first stage) Rotor tip/stator hub clearance Flow coefficient Load coefficient Reaction

characteristics mentioned above, Case 1, and the other is a blade with an increased inlet blade angle around end-wall region, Case 2. The two cases of blade are compared in Fig. 5.36. A blade profile was selected for Case 2 so that thickness of endwall boundary layer was reduced in comparison with that of Case 1. Fig. 5.37 shows operation in progress of the 15 kg/s helium flow, 3.65 MW motor-driven test rig for the model compressor. The test compressor inlet pressure is adjustable up to 1.0 MPa by varying helium inventory in the closed circuit. The test compressor inlet temperature is controllable by a helium-to-water cooler. The helium circuit includes parallel valves for crude and fine flow throttling to accurately regulate compressor pressure ratio. A computerized data acquisition system provides real-time measurement and data reduction from a total of 65 instrumented stations on the compressor and is put in place of the uppermost industrial quality assurance for the subject test operation, instrumentation calibration, and measurement. Pressure and temperature wall taps are present at all interstage locations for monitoring performance health of individual stages. In order to investigate the flow behavior in the blade row of the compressor model in detail, numerical analysis was conducted using the CFD code. The CFD code employed is a fully three-dimensional multistage viscous solver extensively used in-house for production gas turbine development. It is based on such

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Case 1

Case 2

(A)

363

Case 1

Case 2

(B)

Figure 5.36 Comparison of blade shapes between Cases 1 and 2 [30]. (A) Rotor blade. (B) Stator blade.

Figure 5.37 Helium compressor test rig [7].

application experience that the code calibration is focused on mesh refinement, selection of turbulence model parameters, and establishment of appropriate computational boundary conditions in benchmark against available test data. Table 5.16 shows the results of performance test and numerical analysis between the inlet and outlet of the cascade at the design point of the compressor model. The experimental results of the adiabatic efficiency were 88.4% in Case 1 and 87.7% in Case 2, which were in good agreement with the numerical results. Also, the numerical results of the pressure ratio were in agreement with the experimental results within an accuracy of about 2%.

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Table 5.16 Comparison of experimental and numerical results of performance of compressor at design point [30]. Case 1

Pressure ratio Adiabatic efficiency (%)

Experiment

Analysis

Experiment

Analysis

1.168 88.4

1.168 87.4

1.163 87.7

1.143 87.5



Experiment (Case 1) Numerical analysis (Case 1) Numerical analysis (Case 2)

100

100

80

80

80

60 40 20

Inlet Guide vane

0 50 100 150 200 Axial velocity (m/s)

Height (m)

100

Height (m)

Height (%)

Case 2

60 40 20

2nd Rotor blade

0 50 100 150 200 Axial velocity (m/s)

60 40 20

4th Rotor blade

0 50 100 150 200 Axial velocity (m/s)

Figure 5.38 Comparisons of experimental and numerical results of axial velocity between Case 1 and Case 2 [30].

Fig. 5.38 shows the distribution of axial flow velocity in the blade height direction at inlet of the inlet guide vane, and outlet of the second- and fourth-stage blades in Case 1. The symbols indicate the experimental results of Case 1, and the solid and dotted lines indicate the numerical results of Cases 1 and 2, respectively. The distribution of the experimental axial flow velocities in the blade height direction has hardly changed after the second stage, and the flow through the repeating stage was confirmed from this result. Significant decrease in axial velocity was not observed in the main region with a blade height of 10%85%. However, in the fourth stage of the rotor blades, the axial flow velocity decreased to around 84 m/s near the blade tip with a blade height of 95%. The decrease is considered to be caused by the leakage flow through the gap between the blade tip and the casing. On the other hand, the axial flow velocity at the base of the blade with a blade height of 5% was about 127 m/s, which was high as compared to the blade tip. This is because there is no gap at the base of the rotor blade. The numerical results were well predicted to experimental values, and the validity of the analysis could be confirmed.

0.06 1–η∼Re–0.35

0.08

Re extraporation (full scale)

0.10

Case 1 Case 2

0.14 0.16

88 84

Recritical=1.2×106 for full scale

1.00E+06 Reynolds number

90 86

Recritical=4×105 for 1/3 scale

0.20 1.00E+05

94 92

Throughflow prediction (full scale 20 stagers)

0.12

0.18

365

82

Polytropic efficiency

Polytropic inefficiency 1-η

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80 1.00E+07

Figure 5.39 Reynolds number dependency of compressor efficiency [30].

The relationship between the polytrophic inefficiency, 1 2 η, in the blade cascade of the test compressor and the Reynolds number is shown in Fig. 5.39. The Reynolds number is defined by the following equation. Re 5

cV ; ν

(5.5)

where c, V, and ν are chord length, mean velocity, and viscosity of helium gas, respectively. Although the polytrophic inefficiency decreases with the increase in the Reynolds number, there is almost no change in the polytrophic inefficiency in the region where the Reynolds number is higher than about 4.5 3 105. In the case of an axial compressor, assuming that the polytrophic inefficiency is inversely proportional to the power of the Reynolds number, the relationship between the efficiency of the full compressor and the test model was given by the following equation.  n 1 2 ηf Ret 5 ; 1 2 ηt Ref

(5.6)

where n is the power number, and the subscripts f and t are full and test, respectively. From the approximate curve of the experimental results, the power number, n, has the following numerical value.  n5

0:35 Re , Recritical : 0 Re , Recritical

(5.7)

It is known that the polytrophic inefficiency does not decrease in the region above the critical Reynolds number, Recritical, due to the surface roughness of the blade. In order to calculate the critical Reynolds number of the full compressor, the Reynolds number is defined using the surface roughness of the blade, kcla, as

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a characteristic linear length. In order to apply the experimental equation of flow resistance to surface roughness [33], Eq. (5.5) was modified as follows. Rek 5 kcla

V Recritical 5 kcla ; ν c

(5.8)

where the average value of the surface roughness measured along the center line in the blade chord direction on the negative-pressure surface of the rotor and stator blades was used as kcla. The critical Reynolds number of the full compressor is calculated to be 1.2 3 106 based on that of the test compressor, 4.5 3 105. As shown in Fig. 5.38, the polytrophic efficiency of the full compressor is estimated to be about 92% by extrapolating the polytrophic inefficiency to the critical Reynolds number of the full compressor. Based on this result, the adiabatic efficiency of the full compressor considering the losses in the suction and discharge casings is estimated to be about 89.7%. It was confirmed that the design target value of the full compressor of 89% [34] can be achieved.

5.5.3 Magnetic bearing Main specifications and outline of gas-turbine rotor of HTGR are shown in Table 5.17. The generator is a helium gas-cooled synchronous type generator with the maximum power output of 310 MVA. The unbalance grade of the generator and turbo-compressor was set to G2.5 in the ISO1940/1. The calculation of unbalance responses of the rotors confirmed that the vibration amplitudes were well below the allowable limits based on the standard ISO 7919-3 at the rated speed. The allowable vibration amplitudes at the rated speed were 75, 125, and 200 μmp-p (peak to peak) for levels of normal, alarm, and trip, respectively. The allowable vibration amplitude at critical speeds was set at 125 μmp-p. The flow diagram of MIMO control system is shown in Fig. 5.40. The controller is composed of a 4-input 4-output HN controller and 4-phase compensators each of which is connected to each output of the HN controller. The phase compensator is composed of a proportional, phase shift, and phase delay compensators for loop shaping transfer function [35] of a control object, which consists of a rotor and magnetic bearings. This is because the HN controller cannot do loop shaping by itself. On the other hand, the HN controller carries out MIMO control, which links magnetic forces of multiple magnetic bearings, which cannot be realized by phase compensators only. The magnetic bearing consists of a sensor for measuring vibration displacement of rotor, an electric magnet, an electric amplifier, which supplies electric current based on a command from the controller, and a current feedback. The MIMO controller lowers the deviation between set and measured values to control the vibration at the bearing, which uses four feedback vibration measurement values. The sensor is an inductance type for measuring vibration displacement. The numerical model of equivalent mass distribution of rotor is shown in Fig. 5.41. The number of discretization for axial direction is 176. The numerical

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Table 5.17 Main specifications of actual magnetic bearing suspended rotor system [35]. Turbo-compressor rotor

Generator rotor

Coupling type Orientation

Flexible type Horizontal

Mass (kg) Total length (mm) Bearing span length (mm)

51,784 12,900 10,690

66,517 12,800 8190

Inner diameter (mm) Length (mm)

750 740

Static load (t)

25.3 (turbine side) 20.8 (compressor side) 6 6.3 (turbine side) 6 5.2 (compressor side)

900 790 (free side) 820 (turbine side) 32.1 (free side) 34.4 (turbine side) 6 8.0 (free side) 6 8.6 (turbine side)

50 6 2.5

10 6 2.5

Radial bearing

Dynamic load (t)

Thrust bearing Static load (t) Dynamic load (t) Rotor balance quality Response amplitude

ISO 1940 balance quality grade G2.5 # 75 μm peak to peak at rated speed # 125 μm peak to peak at critical speed Rated speed 3600 rpm 0%110% of rated speed condition in normal operation

Speed range

+

Compensator

-

Amplifier

Magnet

Sensor

X

Current feedback

Compensator H∞ controller

+

Amplifier

Magnet

Current feedback

Compensator

Amplifier

Magnet

Sensor Rotor

+

Sensor

-

Y

X

Current feedback

+ -

Compensator

Amplifier

Magnet

Sensor

Current feedback

Figure 5.40 Flow diagram of multiinput multioutput control system [32].

Y

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High Temperature Gas-cooled Reactors

Figure 5.41 Equivalent mass distribution of numerical model [32].

Normalized displacement

1.0

0.0

1.0

–1.0 Normalized distance from the end of exciter

Figure 5.42 Numerical result of vibration mode of rotor at the fourth critical speed for bearing stiffness 109 N/m [32].

result of vibration mode at the fourth critical speed with a bearing stiffness of 109 N/m is shown in Fig. 5.42. The horizontal axis is the distance from the end of exciter normalized by the total length of rotor. The vertical axis is the vibration amplitude normalized by its maximum value. The bearing positions are depicted by Δ. The numerical result shows that an inclination of the central part of generator rotor where mass concentrated was significant, and that the vibration amplitude is higher in the exciter side than that in the coupling side. The unbalance mass was set by exciting the second vibration mode in which the maximum vibration amplitude of rotor occurred in previous calculations [34]. Such a vibration mode in which the central part of the generator has a large inclination occurs at the fourth critical speed. In the analysis, two unbalance masses were set at the both ends of the inclination of the central part of rotor as shown in Fig. 5.43. The unbalance positions were selected so that the maximum vibration amplitude occurs at the fourth critical speed as shown in Fig. 5.43. To investigate the effect of reduction methods, the vibration amplitude of unbalance response, which depends on rotational speed, was compared between the controller transfer functions, which were derived using the conventional modal decomposition method and the modal decomposition using Schur decomposition. The transfer functions were derived at the fourth critical speed at which modeling and

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Exciter

U=2.21 N/m Core Coupling

U=2.21 N/m Magnetic bearing

Magnetic bearing

Figure 5.43 Unbalance mass position [32].

140

Vibration amplitude (μmp-p)

Allowable limit at critical speeds: 125 μmp-p

120

Exciter side Conventional modal decomposition Modal decomposition using Schur decomposition

100

p-p 80 Allowable limit at rated speed: 75 μm

60

Coupling side Conventional modal decomposition Modal decomposition using Schur decomposition

40 20 0 0

600

1200

1800

2400

3000

3600

4200

Rotational speed (rpm)

Figure 5.44 Effect of reduction method on unbalance response [32].

reduction errors are considered sufficiently small. The rotational speed was varied up to 110% of the rated rotational speed. The numerical result of the vibration amplitude of unbalance response of rotor is shown in Fig. 5.44. The horizontal axis is a rotational speed and the vertical one is vibration amplitude of rotor. The numerical result using the modal decomposition using Schur decomposition is shown by a solid line, and the conventional modal decomposition method is shown by a dashed line. The results of exciter and coupling sides are shown by bold and thin lines, respectively. As shown in the mode diagram by numerical result of natural vibration in Fig. 5.42, the vibration amplitude of exciter side was prospected to be higher than that of coupling side, as well as to be maximum value at the fourth critical speed, because the unbalance mass position was set so that the forth vibration mode should be excited. The numerical result of unbalance response predicted adequately that the vibration amplitude of the exciter side was the maximum at the fourth critical speed around 2400 rpm. By using the modal decomposition using Schur decomposition, the vibration amplitude was lowered below 53 and 80 μmp-p at the rated rotational speed and the fourth critical speed, which were lower than the allowable limits of 75 and 125 μmp-p, respectively. The vibration amplitude was reduced to 32 and 37 μmp-p, at the rated and critical speeds, respectively, by MIMO controller, in

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High Temperature Gas-cooled Reactors

comparison with 63 and 78 μmp-p by single-input single-output (SISO) controller, which was almost close to the allowable limit [34]. The numerical results adequately evaluated that the vibration amplitude can be restricted below the allowable limits by deriving the controller transfer function at the fourth critical speed, in which a modeling error was prospected sufficiently low. By using the modal decomposition method, the numerical result of vibration amplitude was 83 and 110 μmp-p at the rated rotational speed and the fourth critical speed. The vibration amplitude at the rated rotational speed exceeded the allowable limit. In comparison with the modal decomposition method, it was shown that the reduction error principally can be reduced by using the modal decomposition using Schur decomposition. The numerical analysis adequately evaluated that the controller was stably derived to reduce the vibration amplitude of rotor below the allowable limits by lowering the reduction error by the modal decomposition using the Schur decomposition. Based on the above analytical research, the conceptual design of the control system for magnetic bearing has been completed. The demonstration test is left as one of the future development issues.

5.6

Iodinesulfur process technology for hydrogen production

5.6.1 Introduction Hydrogen (H2) energy has been attracting attentions as a candidate source of energy system that is sustainable in resources. To create a society whose energy supply comes from H2, it is essential to have large-scale primary energy and resources. Technologies are also necessary that are environmentally friendly and economical and that are suitable for producing H2 in considerable volumes. To meet all these requirements, HTGR and the thermochemical H2 production process have been attracting attention in recent years. Thermochemical H2 production is a method of water splitting to produce H2 through a few chemical reactions. Over 100 kinds of processes using various compounds have been proposed. Iodinesulfur (IS) process is considered the most promising one and this process has been most deeply investigated. One of the expected heat source is hightemperature thermal energy of 950 C obtained from HTGRs. The IS process has a potential for large-volume H2 production at high efficiency. The IS process consists of the following reactions, which proceed thermally: SO2 1 I2 1 2H2 O ! H2 SO4 1 2HI;

(5.9)

H2 SO4 ! H2 O 1 SO2 1 0:5O2 ;

(5.10)

2HI ! H2 1 I2 :

(5.11)

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In the Bunsen reaction, Eq. (5.9), water (H2O) as a raw material reacts with iodine (I2) and sulfur dioxide (SO2) to produce hydrogen iodide (HI) and sulfuric acid (H2SO4). The Bunsen reaction proceeds exothermically below 100 C. Endothermic H2SO4 decomposition, Eq. (5.10) at 900 C produces oxygen (O2); endothermic HI decomposition, Eq. (5.11) at 400 C produces H2. Many research institutes have/had studied the IS process [36]. For example, General Atomics, Sandia National Laboratories, USA and the French Alternative Energies and Atomic Energy Commission collaborated to conduct an integrated laboratory-scale experiment mainly in 2000s. In Korea, a test facility with the H2 production capacity of 50 L/h was constructed using industrial structural materials and was operated since 2010. In China, a test facility made of glass and fluoroplastic was operated for several hours at the H2 production rate of 10 L/h in 2009. A bench-scale experiment was finally completed with the H2 production rate of 60 L/h in 2014. JAEA has also been conducting R&D since 1980s with operation of IS process test facilities by taking several steps. In one step, a facility of the total IS process was constructed and operations were demonstrated. When the tests were successful, the R&D progressed to the next step with a larger scale facility.

5.6.2 Bench-scale test After a lab-scale test as the first step to confirm continuous closed cycle operation of the IS process for 48 h, whose H2 production rate was 0.001 m3/h, a bench-scale test was carried out. There were practical difficulties in conducting such operations: handling corrosive acids, loss of chemicals in side reactions and discharge of SO2, clogging of pipes due to solidification of I2, liquidliquid separation of Bunsen reaction product, and maintaining production rate of the Bunsen reaction. To demonstrate that the operating demands can be satisfied, H2 production operations were performed using an experimental facility [37]. The facility is made of fluorine resin, glass, and quartz; it operates at atmospheric pressure. They are fixed to a steel frame, which is about 5 m wide, 4 m deep, and 5 m high. The facility is equipped with flow rate controllers, liquid-level controllers, and temperature controllers. The heat required for operation is supplied by electricity. The experiment lasted for 20 h in the closed cycle. Temperatures and liquid levels of each equipment were almost constant. In addition, no SO2 was detected at the top of the Bunsen reactor. Fig. 5.45 shows cumulative amounts of H2 and O2 production; H2 production is almost stable at the rate of 31.5 L/h, and the production ratio of O2 to H2 was almost 0.5:1. The experiment span was longer than the time of replacement of the entire HI and H2SO4 inside the facility [37].

5.6.3 Elemental technologies Components in the bench-scale test were mainly made of glass. Components made of industrial materials such as metal and ceramics are needed in an industrialized IS plant because glass material is not proper for scale-up. Environment of the IS

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High Temperature Gas-cooled Reactors

Figure 5.45 Cumulative H2 and O2 production in bench-scale test [37].

process is very severe: high temperature, high pressure, and high corrosive chemicals. Therefore special components to resist such environment were prototyped. H2SO4 vaporizes in a H2SO4 decomposer using high temperature helium gas. Silicon carbide (SiC) ceramic heat exchanger was proposed considering its high corrosion resistance against H2SO4 of liquid/gas phases. At first, a block decomposer made of SiC was prototyped [38]. This decomposer was composed of two stacked circular blocks of 0.75 m in height and has multi-hole channels separated into helium and H2SO4 flows for heat exchange. The heat transfer simulation showed that the maximum thermal stress due to heat exchange was as large as about 126 MPa which was about half of the average tensile strength of SiC material (249 MPa). As described in the next section, therefore, a tube-type (bayonet) decomposer that could reduce thermal stress by one order was manufactured. A pump for H2SO4 transportation was prototyped [39] because there was no pump to transport high temperature and high pressure H2SO4. SiC ceramic is applied for the parts contacting H2SO4 directly. Test operation showed over 90% of volumetric efficiency (the ratio of measured flow rate and designed flow rate) was resulted for H2O and H2SO4 at 100 C. Satisfactory transportation property of the pump in the condition of high temperature and high concentration H2SO4 is expected by the innovation.

5.6.4 Industrial material component test After the success of the bench-scale test, JAEA progressed its R&D step to industrial material component test. Industrial materials with corrosion and thermal resistance were used instead of glass because glass material was not proper for scale-up. JAEA confirmed basic functions and controllability of main components, namely, the Bunsen reactor, the H2SO4 decomposer, the electro-electrodialysis (EED) stack, and the HI decomposer made of such industrial materials before a test of total IS process. The Bunsen reactor is an external-circulating gas-liquid co-current-type reactor [40]. A test was conducted to verify main functions of the reactor. Multiple

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Figure 5.46 External-circulation gasliquid co-current-type Bunsen reactor [36].

functions were expected by combining a static mixer for the reaction, pump for circulation, heat exchanger for heat removal, and reservoir tank for gasliquid separation, as shown in Fig. 5.46 [36]. HIx solution (HI, I2 and H2O) was circulated in the circulation line. SO2 gas, HIx solution, H2SO4 solution, H2O, and nitrogen gas as carrier gas were simultaneously fed into the line. Operating temperature was 20 C and operation pressure was atmospheric pressure. The material flow rate was equivalent to 20 NL/h-H2. As a result, the outlet gas flow rate was consistent with the nitrogen gas flow rate. SO2 gas fed as the reactant was completely absorbed into the Bunsen reaction solution. The main functions of the Bunsen reaction were verified. A bayonet-type heat exchanger composed of SiC ceramics was used as the H2SO4 decomposer. The functions of H2SO4 vaporization, sulfur trioxide (SO3) decomposition, and internal heat recovery were integrated into a unit. Fluid leakage at seal parts and thermal stress due to thermal expansion during high temperature service can be eliminated. A manifold made of glass-lined steel was used for stable operation at high temperatures, as shown in Fig. 5.47. The H2SO4 decomposition section test was conducted to confirm the performance of the decomposer [41]. The H2SO4 decomposer was designed to produce O2 at the rate of 50 NL/h (equivalent to the H2 production rate of 100 NL/h). The test pressure was approximately atmospheric pressure, and the maximum temperature was 850 C. Approximately 90 wt.% H2SO4 solution was fed into the decomposer. The solution was heated and vaporized. H2SO4 was decomposed on a 0.1 wt.% Pt/TiO2 catalyst bed. O2 gas was separated from the decomposition product by cooling and an active carbon absorber. Fig. 5.48 shows that the O2 production rate is proportional to the H2SO4 feed rate. This result is

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High Temperature Gas-cooled Reactors

O2 production rate (NL/h)

Figure 5.47 Bayonet-type H2SO4 decomposer [41].

100 80 60 40 20

H2SO4 feed: 90 wt.% Maximum temp: 850°C

0 0

5

10

H2SO4 feed (mL/min)

Figure 5.48 O2 production rate against H2SO4 feed rate to decomposer.

evidence that the decomposer achieved the design performance and the O2 production rate can be controlled easily by adjusting the H2SO4 feed rate. The H2SO4 decomposition ratio of 80 mol.% was estimated based on the experiment result. This value was close to the equilibrium value of 84.4 mol.% at 850 C [42]. The result confirmed satisfactory performance of the H2SO4 decomposer.

R&D on commercial high temperature gas-cooled reactor

13.0

375

(A) Catholyte

11.0

HI

Molality (mol/kg)

9.0

I2

7.0 5.0 11.0

HI

9.0 7.0 5.0

I2 (B) Anolyte

0

10

20 Current (A)

30

40

Figure 5.49 Concentration change in (A) catholyte and (B) anolyte in outlet solution of EED stack (Plot: experimental value, Line: prediction curve).

The EED is applied to concentrate HI in the HIx solution from the Bunsen reaction section for higher H2 production efficiency of the IS process. The EED stack was developed based on a single-cell study [40]. The EED stack was composed of electrodes made of impervious graphite and a cation-exchange membrane made of Nafion 117 (DuPont). The effective membrane area was 300 cm2 per one layer. Eleven cells were layered in one stack. The anolyte and catholyte solutions flowed along both sides of the membrane. The HI concentration test was conducted using the EED stack to evaluate the performance of the stack [40]. HIx solutions (HI, I2 approximately 10 mol/kg-H2O) were fed to an anode flow channel and a cathode flow channel in the stack at 20 C. Compositions of the outlet solutions of the anolyte and the catholyte were evaluated using a chemical titration method. Fig. 5.49 shows changes in the outlet solution concentrations in different currents. The experimental results agreed with the prediction curves calculated based on the results of a previous study. This agreement demonstrated that the EED stack performed well in terms of concentrating HI. Fig. 5.50 shows the HI decomposer, which is a radial-flow-type adiabatic reactor with a fixed catalyst bed. HI is fed to the outer tube side; then it is decomposed as flowing through the catalyst bed. The decomposed gas of HI, H2, and I2 flows out from the inner tube. It is expected that H2 embrittlement of the outer tube, which works as pressure vessel, is prevented because contact between the produced H2 and the metallic material of the tube is avoided [42]. The HI decomposer was designed to produce H2 at the rate of 100 NL/h. The HI decomposition test was conducted to confirm the performance of the decomposer. Pure HI gas from a gas cylinder was fed into the HI decomposer. The test pressure was approximately atmospheric pressure, and the maximum temperature was 500 C. The flow rate and H2 concentration of the product gas were measured using a gas flow meter and a gas chromatograph, respectively.

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High Temperature Gas-cooled Reactors

Figure 5.50 Radial-flow-type adiabatic fix-bed HI decomposer.

(A)

(B) 30

10

Decomposition rate (%)

20

Equilibrium =22%

Stable

0

Figure 5.51 H2 production performance of HI decomposer: (A) H2 production and HI decomposition rates against HI gas feeding, (B) time-dependent H2 production rate and estimated HI decomposition ratio.

Fig. 5.51 shows H2 production performance of the HI decomposer. The result shown by (A) is evidence that the H2 production rate can be controlled by adjusting the HI gas feed rate. The designed H2 production rate of 100 NL/h could not be achieved because a large amount of HI gas could not be fed into the decomposer owing to cooling of the feed gas cylinder. The result shown by (B) indicates the time dependence of the H2 production rate and the estimated HI decomposition ratio [42]. The H2 production rate of c. 20 NL/h and the HI decomposition ratio of c. 18 mol.% were achieved in a stable manner over 3 h of operation. The HI decomposer performed well in terms of H2 production.

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5.6.5 Hydrogen production test Construction of the H2 production test facility was completed in March 2014 to verify the integrity of all IS process components made of industrial structural materials and to demonstrate stable H2 production. Table 5.18 lists the specifications of the facility. Fig. 5.52 shows a schematic configuration diagram of the facility. The facility consists of three chemical processing sections: each section was designed to be operated

Table 5.18 Specifications of H2 production test facility [36].

Size of facility Width (m) Depth (m) Height (m) Ability of H2 production rate (L/h) Designed maximum temperature (K)

18.5 5.0 8.1 100 1123.15

Representative structural materials SiC ceramics Hastelloy C-276 Impervious graphite Glass-lined steel Fluororesin-lined steel

H2SO4 decomposer HI decomposer and pipes EED stack and heat exchanger Pipes and tanks Pipes, valves and tanks

Figure 5.52 Schematic diagram of H2 production test facility.

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High Temperature Gas-cooled Reactors

individually for shakedown operation. The designed H2 production rate was 100 NL/h. The heat was supplied from electric heaters [36]. The EED stacks are first applied to the test facility in the step. Main components, namely, the Bunsen reactor, the H2SO4 decomposer, the EED stacks, and the HI decomposer, whose functions were confirmed as shown in the last section, were applied in the facility. As for the first test, conducted in February 2016, JAEA succeeded continuous H2 production of 10 L/h-H2 for 8 h with integration of the three sections. Fig. 5.53 shows H2 and O2 production amounts. O2 production was estimated from the H2SO4 decomposition rate calculated from composition data of undecomposed H2SO4. The production ratio of H2 and O2 was around 2, the stoichiometric ratio of H2O decomposition [43].

Figure 5.53 Results of hydrogen production tests [43]. (A) First test of 10 L/h-H2 for 8 hours. (B) Second test of 20 L/h-H2 for 31 hours.  O2 production rate was estimated from H2SO4 decomposition rate.

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In the first test, the pump used to feed HI solution with high I2 concentration malfunctioned, leading to abnormal vibration and abnormal stoppage. Precipitation of I2, which was dissolved in HI solution into the narrow gap between a cylinder and a piston, caused the malfunction. To prevent such I2 precipitation, a shaft seal technology was developed using both nitrogen gas and HI solution injection into the narrow gap of the shaft seal part to flush out solid I2. Nitrogen gas was fed continuously to purge the HIx solution and the vapor from the gap. HI solution injection is effective for solution and removal of the precipitated solid I2 in the gap. In addition, vapor evacuation above the piston and the cylinder is made to prevent a small amount of leakage from the top of the gap. The pump equipped with the developed shaft seal technology was verified under environmental conditions similar to those when the pump stopped functioning. The modified pump worked continuously and stably for 24 h. Effectiveness of the developed technology was confirmed [36]. The second test was conducted in October 2016 after application of the shaft seal technology. The technology was confirmed successfully, and H2 production rate and test duration were increased to 20 NL/h-H2, 31 h as shown in Fig. 5.53B [43]. I2 plugging, however, occurred in the heat transfer tubes of the cooler in the Bunsen reaction section. Increase of H2O content in the Bunsen reactor solution led to a decrease in the I2 solubility of the solution. I2 was precipitated easily in the heat transfer tubes of cooler, a point of low temperature. At present, JAEA is investigating the cause for the increased H2O content of the solution and a control method of composition fluctuation by controlling heating and flow rate of chemical reactors to achieve stable and longer term operation [36]. After the second test, conditions of corrosion-resistant component material were investigated by overhaul inspection. Though serious damage or corrosion was not found, there was slight damage to some components. JAEA maintained the components and made a maintenance plan of the facility [43].

5.6.6 Improvement of hydrogen production efficiency In addition to the above-mentioned industrial material component test and H2 production test, JAEA has been making R&D on improvement of H2 production efficiency of the IS process. The R&D is required for commercialization of the process. When the efficiency is high, heat and electricity requirement to produce a certain amount of H2 are smaller. H2 production cost can be reduced by the smaller demand.

5.6.6.1 Electro-electrodialysis cell When HIx solution from the Bunsen reaction section was provided directly to a HI distiller, about 5 mol of H2O was expected to be vaporized with 1 mol of HI by distillation due to pseudo-azeotropy of the solution. The H2 production thermal efficiency was estimated low because excess heat to vaporize H2O was required. EED has been proposed to avoid the problem by concentration of HI to over azeotropy.

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High Temperature Gas-cooled Reactors

JAEA has studied on performance parameters of the EED cells because the performance has a great influence on the electricity demand of the EED cells and heat demand of the HI distiller. Amounts of HI and H2O increase and the amount of I2 decreases in the HI-rich catholyte through electrochemical reactions at the surface of the electrodes as in Eqs. (5.12) and (5.13) and permeation of H1 with H2O through a cation exchange membrane. Material balance is determined by the parameters of proton transport number (t1) and electroosmosis coefficient (β). H1 of 2t1 mol and H2O of 2t1β mol transfer from anolyte to catholyte through a cation exchange membrane per I2 production of 1 mol on anode. In addition, I2 of 2 (1 2 t1) mol transfers from catholyte to anolyte. 2I2 ! I2 1 2e2 ðAnodeÞ;

(5.12)

I2 1 2e2 ! 2I2 ðCathodeÞ:

(5.13)

Performance of the cell is determined by parameters of t1 and β, and initial cell voltage (Einit.). High value of t1 and low value of β are advantageous. Small Einit. has an advantage because electricity consumption is smaller. Influence of temperature and solution composition on the performance parameters was evaluated. EED tests were carried out applying a commercial Nafion 112 membrane (DuPont) and ETFE-St. ETFE-St was prepared in JAEA by grafting styrene (St, 99%, Sigma Aldrich Japan K.K.) on (poly(ethylene-co-tetrafluoroethylene)) (Asahi Glass Co.) membrane. Initial solution concentrations of HI and I2 in the anolyte and catholyte cells were set as follows: mHI 5 7, 10, 12, and 13 mol/kg (mHI =mI2 5 1) or mHI =mI2 5 0:5, 1, and 2 (mHI 5 10 mol/kg). Here, mi means molality of component i. Test temperature was 40, 70, and 100 C, respectively. The fitting parameters of equations to determine the performance parameters were regressed by least square method applying the test results. Parameter set from the fitting parameters reproduced the measured result with 6 20% in both the membranes cases [41]. These equations were used to determine operation conditions of the H2 production test to verify integrity.

5.6.6.2 Hydrogen iodide decomposer with hydrogen separation membrane The efficiency-determining step in the IS process is the HI decomposition section owing to low conversion of HI decomposition into H2 at chemical equilibrium (0.22 at 400 C). This low conversion increases the amount of recycled materials in the HI decomposition section, thereby increasing the thermal burden and decreasing the thermal efficiency of the process. The development of a membrane-based H2 separation process can open up the possibility of increasing the conversion of HI decomposition. The feasibility of enhancing conversion of HI decomposition using membrane reactors has been studied theoretically and experimentally. Myagmarjav et al. [44] reported a significant enhancement in conversion of HI decomposition (0.60) using

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2.5

Conversion (-)

0.8

2 0.6 1.5 0.4 1 0.2 0

0.5

2

4

6

8

10

H2 flow rate (mL min-1)

3

1

0

HI flow rate (mL min-1)

Figure 5.54 Conversion of HI decomposition and H2 production rate as function of the HI feed flow rate in silica membrane reactor [44].

a membrane reactor equipped with a tubular silica membrane for midtemperature (400 C) for the first time. The membrane reactor comprises two concentric channels. The annulus between the channels is the reaction zone where HI decomposition takes place with the active carbon catalyst. H2 generated by HI decomposition permeates from annulus to the inner tube through the wall of the silica membrane. The permeated H2 is carried away by sweep gas. Fig. 5.54 shows the experimental results of HI conversion and the H2 production rate as a function of the HI feed flow rate [44]. The conversion attained at an HI flow rate of 2.6 mL/min was approximately 0.60, which was greater than the equilibrium conversion in HI decomposition of 0.22 at 400 C. This result indicates that H2 was extracted through the silica membrane, which constituted a successful demonstration of the silica membrane reactor. A significant increase in conversion of HI decomposition was achieved. Moreover, the operating features to control the performance of the membrane reactor in particular, total pressures on the feed and permeate sides, HI decomposition temperature, and HI feed flow rate over a wide range were investigated theoretically by the simulation [45]. The membrane reactor was found to have conversion limitations, and the HI conversion can be improved by up to about four times (0.80) or greater than the equilibrium conversion (0.22) at 400 C by employing the silica membrane reactor. The simulated results showed good agreement with the experimental findings.

5.6.6.3 Analytical estimation of hydrogen production thermal efficiency The H2 production thermal efficiency of the IS process has been evaluated by process flow analysis inside and outside of JAEA. Some literature report thermal efficiency of over 40%. Recently, JAEA conducted a study to achieve higher

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High Temperature Gas-cooled Reactors

efficiency by applying innovative techniques: the HI decomposer with the H2 separation membrane, reverse osmosis (RO) membrane stacks, depressurized flash concentration of H2SO4 by using waste heat from the Bunsen reaction, and directcontact heat exchanger (DCHX) [46]. A process flow was designed by material and heat balance evaluation using PRO/II (Schneider Electric, SA), a commercial chemical process simulator. The H2 production thermal efficiency (η) was estimated by Eq. (5.14). η5

nH2 ΔHH2 3 102 ; Q 1 ðW=ðηel: 3 1022 ÞÞ

(5.14)

where nH2 , ΔHH2 , Q, and W are H2 production rate (mol/s), HHV of H2 (MJ/mol-H2), total heat demand of the IS process (MJ/mol-H2), total electricity input including utilities such as circulator of helium heat source (MJ/mol-H2), respectively. HHV of 0.2858 MJ/mol-H2 (subscript t means heat) was the combustion heat of H2 to H2O in liquid phase. The net electricity generation efficiency (ηel.) of 50.4% was assumed considering maximum potential of electricity generation in HTGR. The process flow was different from the one of the H2 production test facility to verify integrity mentioned in Subsection 5.6.4. In the H2SO4 section, depressurized seven stages of flash drums were applied for H2SO4 solution. H2SO4 from the 7th drum is further concentrated in a DCHX. H2SO4 from the DCHX is vaporized and decomposed into H2O, SO2, and O2 in the H2SO4 decomposer. Undecomposed H2SO4 from the H2SO4 decomposer is fed to the DCHX to be recovered in the solution from the 7th drum. In the HI concentration subsection, RO membrane stacks were added before EED stacks to concentrate HI. The stack is separated into high pressure side and low pressure side by a RO membrane. Some H2O is permeated from the high-pressure side solution to low-pressure side solution by RO. Then HI is concentrated. In the HI decomposition section, the HI decomposers with the H2 separation membrane are applied instead of radial-flow-type one. The product H2 from HI decomposition reaction is permeated through the membrane and is obtained as product of the IS process. Operating parameters were chosen based upon reasonable assumption of future improvements to the system; purification of H2SO4 and HIx solutions obtained from Bunsen reaction without using HTGR heat, ideal material transfer between anode side and cathode side and low voltage in the EED stack, an RO membrane with perfect H2O permselectivity, an HI decomposition reactor with H2 separation membrane, and heat exchangers of small temperature difference. Innovative techniques were applied to the process design; depressurized flash concentration of H2SO4 using waste heat from Bunsen reaction, prevention of H2SO4 vaporization from the DCHX by introduction of H2SO4 solution from the second flash drum, and I2 condensation heat recovery by introduction of HI decomposition by-product into the HI distiller. Table 5.19 summarizes the heat and electricity demand of the analysis [46]. The demands were standardized to heat input of 170 MWt assuming heat supply from one GTHTR300C. The H2SO4 decomposer has the majority of the heat demand. The major electricity demands were for the EED stacks and the circulator of helium

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Table 5.19 Heat and electricity input and H2 production thermal efficiency [46]. Heat Section/subsection

Component

Heat (MWt)

H2SO4

DCHX H2SO4 distillation column reboiler H2SO4 decomposer HI distillation column reboiler HI decomposition preheater Helium circulatora

1.3 125.0 29.7 19.0 2 5.0

HI separation HI decomposition Utility Total heat

170.0

Electricity

EED Pumps Utility

Component

Electricity (MWe)

EED cell stack Process pumps Vacuum pump Cooling towers circulation pumps Cooling towers Freezer (H2 and O2 products cooler, H2SO4 depressurized flash concentration) Helium circulatora

15.3 2.8 0.1 2.3 0.1 1.9

Total electricity

5.0 27.5

H2 production thermal efficiency H2 production (m3/h) H2 production (mol/s) Net electricity generation thermal efficiency, ηel (%) Net heat input including power generation (MWt) H2 production thermal efficiency (%) a

31,863 394.9 50.4 224.6 50.2

Heat generation from circulator electricity input is recovered.

heat source. The table also shows the H2 production rate and the H2 production thermal efficiency. The H2 production rate was about 31,900 Nm3/h. The thermal efficiency of 50.2% was attained. The thermal efficiency increased about 10% with the contribution of the innovative technology and high-performance components.

5.6.7 Component materials Corrosive compounds such as H2SO4, HI, and I2 are used in the IS process. High temperature of c. 900 C and high pressure of c. 2 MPa are in some components. Component materials have to be developed because some components are exposed to severe environment and common materials cannot be applicable.

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5.6.7.1 Corrosion resistance

Average corrosion rate (mm/y)

Glass-lined components generally indicate excellent corrosion resistance for various corrosive solutions including H2SO4 and HI. However, corrosion data of lining glass for HIx are nothing. Therefore corrosion test for HIx was conducted. Test specimens were immersed in test solution. The glass of the test specimens has the same composition as that of the glass-lined sheath. Corrosion rate was determined by weight change of the test specimens before and after corrosion test. Test solution was heated by a mantle heater and a hot stirrer up to approximately 115 C. The composition of test solution (HI:I2: H2O 5 0.9:1.4:7.7) was same as that of HIx purifier in Bunsen reaction section of the H2 production test to verify integrity. Fig. 5.55 shows average corrosion rate. The corrosion rate was around 0.030.04 mm/y at 50 h, and under 0.02 mm/y after 100 h [47]. These values are low enough in comparison with the standard of manufacturer (0.1 mm/y). This is the same trend seen in the previous result for other type of glass in H2SO4. H2SO4 concentration is one of the severest environments for materials. Concentration range is 5090 wt.% and temperature range is 200400 C. Glass lining is considered as one of the candidate component materials for the part. Corrosion resistance and thermal resistance of the material were investigated. The immersion test of a soda-lime glass in H2SO4 was carried out in an autoclave. Fig. 5.56 displays the S-1

S-2

S-3

0.10 0.08 0.06 0.04 0.02 0.00

0

20

40

60

80

100

120

140

160

Time (h)

Figure 5.55 Average corrosion rate of the test specimens up to 150 h [47].

Figure 5.56 Corrosion rate of glass lining in H2SO4 [48].

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corrosion rate [48]. The rate of 0.1 mm/y is considered as the upper limit of allowance. Though some tests showed high corrosion rate in short time immersion, good corrosion resistance around 0.1 mm/y was resulted after the immersion time of 6090 h. Alkali ion removal from surface was assumed as the reason of the initial weight decrease and high corrosion rate in the short time immersion.

5.6.7.2 Strength Ceramic materials have high resistance against high temperature and corrosive compounds. However, fracture occurs in ceramic materials. Since the fracture of ceramic material occurs due to the presence of small cracks or crack-like flaws contained in the material body, which acts as stress concentrators, the variation of strength is caused by variation of flaw distribution depending on the component design and the material fabrication. Such variation of strength is a problem for application of ceramic materials in the IS process. In general, the effective volume theory is one of appropriate theories for estimating the ceramic component strength. JAEA has proposed the novel evaluation method of SiC component strength by combining an effect of flaw distribution with the effective volume theory. The effect of flaw distribution on SiC strength was evaluated using Monte Carlo simulation [49]. In this method, the estimation formula of minimum strength σmin is given by: σmin 5 293 

 0:013 1 ; Ve

(5.15)

and the effective volume Ve is defined as ð Ve 5

σ

σmax

1=m dv;

(5.16)

where σmax, σ, and m are the maximum stress value of the stress distribution in the component, the stress applied to microvolume dv, and the Weibull modulus, respectively. Although the Weibull modulus varies depending on the size of the component, it can be fixed as 8 to calculate the effective volume based on the results of Monte Carlo simulation in the previous study [49]. Fracture tests were conducted to verify an estimation formula of minimum strength in the strength evaluation method [43]. The strength of various volume components was examined using three types of components: (1) straight tube model, (2) small-scale outer tube model, and (3) full-scale outer tube model. Fig. 5.57 shows shape and size of the models. The full-scale outer tube model has the same dimensions to the outer tube of the H2SO4 decomposer in the H2 production test facility to verify integrity. The straight tube and small-scale outer tube models have the same wall thickness and length. The small-scale outer tube model has a flange for sealing at one end. A model was fixed with fixing jigs in a protection barrier. The pressurized system produced hydraulic pressure by a pump up to 100 MPa, and uniform pressure was loaded on the inner surface of the

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High Temperature Gas-cooled Reactors

Figure 5.57 Dimensions of SiC models for fracture test [43].

Figure 5.58 Estimated minimum strength and fracture strength of SiC models [43].

fixed model. The pressure was continuously increased until fracture. The whole body was shattered uniformly in a moment for every model. This fracture behavior suggested that the entire body of the model was stressed up to the limit of destruction, and values of acquired fracture stress were close to the strength value in every model design. Fig. 5.58 shows an estimated minimum strength line and the fracture strength of the each model [43]. The straight tube model (length (L) of 200 mm) made of the same SiC material as the other models has been examined with the same test

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apparatus in the previous study [49]. This figure also shows the results of three bending tests of SiC test specimens, which are made of the same material as the other models. As for the effect of shape on strength, the strength distributions of the straight tube model (L of 400 mm) and the small-scale outer tube model were almost same. This result showed that the fracture strength of the tube type model was significantly related to the strength of the straight part of the model because the volume of this part occupied majority of the model. Strength evaluation of the straight part was effective for the reliability evaluation of the SiC tube shape component of the H2SO4 decomposer. This procedure has an advantage that calculation of the effective volume becomes a simple way because simulation such as the finite element method is not required. As for effects of volume on strength, the estimated minimum strength decreased with the increase of the effective volume because it was assumed that the size and the amount of internal flaws increased with increasing the volume in the strength estimation method [43]. On the other hand, the fracture strength value of the test models did not decrease sufficiently as estimated, even when the effective volume increases in large value. This result suggested that the internal flaw distribution in this large-scale structure was affected slightly by the volume increase. There is a possibility that this method can be used for safer design because this evaluates the strength in conservative side considering the estimated minimum strength line was below all the plots of fracture test results. After further more verification, this method is expected to become a standard method to obtain a license in accordance with the High Pressure Gas Safety Act in Japan.

5.7

System integration technology for connection of heat application system and high temperature gas-cooled reactor

5.7.1 Introduction Nuclear energy has been exclusively utilized for electric power generation, but the direct utilization of nuclear thermal energy is necessary and indispensable so that the energy efficiency can be increased and energy savings can be promoted in the near future. The hydrogen production is one of the key technologies for direct utilization of nuclear thermal energy. The system integration technology for connection of the hydrogen production system to HTGR is one of the key technologies to put hydrogen production with nuclear energy to commercial use, and its research and development was carried out [50]. Fig. 5.59 shows the concept of the HTGR hydrogen production system. The heat generated in the reactor core is exchanged from the primary helium gas to the secondary one with the intermediate heat exchanger (IHX), and the secondary one is transported to the hydrogen production system passing through the hot gas duct.

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High Temperature Gas-cooled Reactors

Hydrogen production system

HTGR

Hot gas duct

Isolation valve

Chemical reactor

Reactor

Isolation region to mitigate explosion and fire effects Intermediate heat exchanger (IHX)

Figure 5.59 Concept of the HTGR hydrogen production system.

Blast Explosion

Reactor

Intermediate heat exchanger Tritium Hydrogen

Primary He gas

Isolation valve

Thermal absorber (Steam generator) High temperature Isolation valve Secondary He gas

Chemical reactor

Raw material

Tritium Hydrogen

Hydrogen

Process gas

Figure 5.60 Flow diagram of HTGR hydrogen production system and research items of system integration technology.

The transported heat is used in the hydrogen production system for the endothermic reaction of hydrogen production. Fig. 5.60 shows the flow diagram of the HTGR hydrogen production system and research items of the system integration technology as follows: 1. Control technology to keep reactor operation against thermal disturbance caused by the hydrogen production system, 2. Estimation of tritium permeation from reactor to hydrogen, 3. Countermeasure against explosion of combustible gas, and 4. Development of a HTIV to separate reactor and hydrogen production systems in accidents.

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5.7.2 Control technology The reactor and the hydrogen production system are connected by the helium gas loop. A chemical reactor causes the temperature fluctuation of the secondary helium gas by the fluctuation of the chemical reaction, which can be occurred at the normal start-up and the shut-down operation as well as malfunction or accident of the hydrogen production system. If the temperature fluctuation is transferred to the reactor, the reactor will be stopped. Therefore the control technology should be developed to mitigate the temperature fluctuation within 10 C at a thermal absorber outlet because the temperature rise above 15 C compared with the normal temperature at the reactor inlet causes the HTTR reactor scram. JAEA proposed to use a SG as the thermal absorber, which is installed downstream the chemical reactor in the secondary helium gas loop [51]. A simulation test with a mock-up test facility was carried out to investigate performance of the SG for mitigation of the temperature fluctuation and transient behavior of the hydrogen production system and to obtain experimental data for verification of a dynamic analysis code [5256]. Fig. 5.61 shows the schematic flow diagram of the test facility. The test facility had an approximate hydrogen production capacity of 120 Nm3/h and simulates key components downstream from the IHX. An electric heater was used as a heat source instead of the reactor in order to heat helium gas up to 880 C (4 MPa) at the chemical reactor inlet, which was the same temperature as the HTTR hydrogen production system (HTTR-H2). The steam reforming process of methane, CH4 1 H2O 5 3H2 1 CO, instead of the IS process is used for hydrogen production of the test facility. Fig. 5.62 shows a schematic view of the SG, which was a kettle type, with 27 heat exchanger tubes (25.4 mm O.D., 4.0 mm thick and 3500 mm length). Helium gas flows inside of heat exchanger Combustion line

Nitrogen feed line LN2 tank

Pump Evaporator Surge tank

LNG tank Pump Evaporator Surge tank

Methane feed line Water tank

Flare stack

Control Vapor valve condenser

GC Gas chromatograph

Radiator

Steam feed line Fan

Steam reformer

Pump

P

Pressure difference control

Steam generator Tank

Filter

Circulator

Electric heater

P Pressure difference gauge

Hot gas duct

Helium gas circulation loop Figure 5.61 Flow diagram of mock-up test facility for development of control technology [53].

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1150

Cross-section AA Steam outlet A

Heat exchanger tube

Helium gas outlet

Helium gas inlet

A Water feed

Water inlet from radiator 3500 4880 (Unit : mm)

Figure 5.62 Schematic view of steam generator of mock-up test facility [54].

tubes and heat of helium gas was transferred to water of which holding quantity at the rated condition was 1.7 m3. And a heat exchanged between helium gas and water in the SG was designed as 135 kW at rated condition. The radiator was installed above the SG to cool a large quantity of steam produced in the SG in case of loss of chemical reaction and circulation of steam and condensed water circulate between the SG and the radiator. In order to simulate the loss of chemical reaction in the hydrogen production system, a feed of methane to the chemical reactor was suspended during the hydrogen production. Fig. 5.63 shows a detailed flow diagram of the test facility. During the test, helium temperature and pressure at the inlet of SR were controlled at each rated condition, 880 C and 4.1 MPa, respectively. Before the methane feed was suspended, flow rates of methane and steam to the chemical reactor, the steam reformer (SR), were controlled at each rates of 12 and 47 g/s by using control valves A and B, respectively, and the water pressure in the SG was controlled at 4.6 MPa by using a control valve C. After the suspension of methane feed, nitrogen of 30 g/s was introduced to the SR to keep a pressure difference between helium and the process gas in the reaction tube of the SR constant. Steam feed to the SR was stopped, and then a large quantity of produced steam was introduced to the radiator. The experimental and numerical results of the system controllability test for the loss of chemical reaction are shown in Fig. 5.64, where the solid and dashed lines show the experimental and numerical results, respectively. A horizontal axis is the time axis indicating an elapsed time from the suspension of methane feed. Before the start of the suspension of methane, the methane and steam flow to the SR and

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Figure 5.63 Detailed flow diagram of cooling system for helium with steam generator and radiator of mock-up test facility [56].

could be controlled at each rated condition, that is, 12 and 47 g/s as shown in Fig. 5.64A. At the moment of 0 h, the methane flow rate was decreased drastically from 12 to 0 g/s. And at the moment of 0.17 h, the steam flow rate to the SR was decreased drastically from 47 to 0 g/s by the closing of the stop valve C installed at the inlet of a steam superheater (see Fig. 5.64). The hydrogen production rate decreased from about 120 to 0m3/h due to the suspension of methane feed as shown in Fig. 5.64B. Fig. 5.64C shows the helium temperature at the inlet and outlet of the SR and the SG. At the moment of 0 h, the SR outlet helium temperature started increasing from 632 C and it reached to 833 C at 1.4 h due to the loss of chemical reaction. And it leads to the increase in the SG inlet helium temperature, and it increased from 548 to 793 C. The SR outlet and the SG inlet helium temperature increased about 200 and 245 C, respectively. However, helium temperature at the SG outlet showed almost stable value, and profile was the same as that of the water pressure and water temperature in the SG as shown in Fig. 5.64D and E. After the suspension of methane feed, the water pressure in the SG increased gradually from 4.61 MPa at 0 h to 4.75 MPa at 0.32 h because of the increasing of the steam generation rate in SG with the increasing of the SG inlet helium temperature. It resulted in the increasing of the water temperature in the SG from 258.7 C to 260.2 C and the increasing of the SG outlet helium temperature from 262.5 C to 264.5 C. As

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High Temperature Gas-cooled Reactors

Figure 5.64 Experimental and numerical results of the system controllability test for loss of chemical reaction using the mock-up test facility [56]. (A) Change of flow rates of steam, CH4 and N2. (B) Change of hydrogen production rate. (C) Change of helium temperatures. (D) Change of steam pressures. (E) Change of water temperatures in SG. (F) Change of steam flow rate at radiator inlet. (G) Change of water temperature at radiator outlet. (H) Change of flowrate of cooling air.

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a result, it was confirmed that the thermal disturbance of the helium temperature caused by the loss of chemical reaction was successfully mitigated by the cooling system with SG and the radiator. The fluctuation range of the helium temperature was from 21.2 C to 12.0 C at the SG outlet, which is within the target range of the HTTR hydrogen production system.

5.7.3 Tritium permeation It is well known that hydrogen isotopes, hydrogen, deuterium, and tritium, permeate through solid metals. Tritium produced in the reactor core tends to permeate through heat transfer tubes of the IHX and the reaction tubes of the chemical reactor. Further, it is probable that the tritium will mix with the product hydrogen. The target of tritium permeation amount into hydrogen is 11.8 Bq/g-H2. This value guarantees exposure rate of the hydrogen user less than 10 μSv/y, which is less than 1/100 of exposure rate by the natural radioactivity. Therefore estimation of tritium permeation should be done, and a countermeasure to reduce tritium permeation, if necessary. A component test was carried out to investigate the permeability of the material of the IHX tubes, Hastelloy XR, and permeation phenomenon of the reaction tubes, where hydrogen and tritium permeate in the counter direction as shown in Fig. 5.60 [57,58]. An experimental apparatus was composed of a test section consisting of a test tube (inside tube), a measuring tube (outside tube), and an electric heater, a gas supply system, a concentration measurement system, and so on as shown in Fig. 5.65. Hydrogen (H2) and deuterium (D2) instead of tritium were flowed with helium and argon gases. The permeation rate of hydrogen isotopes is in inverse proportion to square pffiffiffi root of the atomic mass 13, that is, the permeation rate of tritium becomes 1= 3 of that of hydrogen. We confirmed that the correlation between the permeation rate and the atomic mass was valid using hydrogen and deuterium. Therefore the permeation rate of tritium can be obtained from the experimental data on hydrogen permeation. The permeation amount, Q [m3(STP)/s], and the permeability, Kp [m3(STP) 21 21 m s Pa20.5], are shown by the following equations: Q5

2π  L  Kp pffiffiffiffi pffiffiffiffiffi   pi 2 po ; ln ro =ri 

Kp 5 Fo exp

 2 Eo ; RT

(5.17)

(5.18)

where L, ri, ro, pi, and po are the test pipe length and inner and outer radius of the test tube; H2 and/or D2 are partial pressures at the test and the measuring tube, respectively; and Fo, Eo, R, and T are a preexponential factor, an activation energy, a gas constant, and a temperature, respectively.

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High Temperature Gas-cooled Reactors

Preheater (1kW)

Measurment system -Temperature -Pressure -Flow rate

Automatic control system

Test tube Measuring tube Heating system Main heater (6 kW) Exhaust system

Measurment system -Hydrogen Cooling system Flow control

Flow control Molecular sieve Gas supply system H2/He,D2/Ar, etc.

Molecular sieve Vacuum system Gas purge system Ar, He, etc.

Figure 5.65 Experimental apparatus for estimation of tritium permeation [58].

The permeability of Hastelloy XR was obtained as follows [57]: 1. Hydrogen Average temperature of the pipe: 8731123 K Hydrogen partial pressure: 1.06 3 102 to 3.95 3 103 Pa Activation energy: Eo 5 67.2 6 1.2 kJ/mol Preexponential factor: Fo 5 (1.0 6 0.22) 3 1028 m3(STP) m21 s21 Pa20.5. 2. Deuterium Average temperature of the pipe: 9731123 K Hydrogen partial pressure: 9.89 3 102 to 4.04 3 103 Pa Activation energy: Eo 5 76.6 6 0.5 kJ/mol Preexponential factor: Fo 5 (2.52 6 0.25) 3 1028 m3(STP) m21 s21 Pa20.5.

Experiment and analysis on counterpermeation of D2 and H2 were performed to investigate the effect of the existence of high pressure H2 on the amount of permeated D2, instead of tritium, when the rate-limiting step of the permeation becomes the diffusion process in metals [8]. Fig. 5.66 shows the experimental (symbols) and numerical (lines) results on the counter-permeation. When D permeates from the inside of the test tube and H permeates from the outside, the amount of permeated D2 through the test tube depends not only on the partial pressure of D2 but also on that of H2 existing outside the test tube. When the partial pressure of D2 in the test tube is lower than 102 Pa and the one of H2 outside the test tube is higher than 10 kPa, the amount of permeated D2 in counterpermeation decreases compared with

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1.4 Without fD (fD=1)

Ratio of permeated D2 (D2-counter/D2-ordinary)

1.2 1.0 Without HD molecules 0.8 Numerical results with fD and HD 0.6 D2 partial pressure 0.4 102Pa 0.2

56Pa

0.0 100

101

Average temperature = 819°C 102

103

104

105

106

H2 partial pressure in the measuring tube (Pa)

Figure 5.66 Experimental and numerical results of permeated deuterium on counterpermeation of deuterium and hydrogen [58].

the one in ordinary permeation. This is because the interstitial sites where D atoms can be dissolved on the metal surface decrease by occupation by a large amount of dissolved H atoms on the surface. On the other hand, when the partial pressure of D2 and H2 is lower than 1 kPa, the interstitial sites do not decrease so much because the amount of dissolved H atoms decreases. In this case, the partial pressure of D2 and HD molecules becomes almost equal as taking account of the equilibrium state of D2, H2 and HD molecules on the surface. Consequently, the amount of permeated D2 on counterpermeation increases relative to the one for ordinary permeation because the amount of dissolved D atoms increases. The amount of permeated D2 on counterpermeation, QD [m3(STP)/s], can be predicted by the following in consideration of occupation ratio of the interstitial sites, at which D atoms can be dissolved on the metal: QD 5

 2π  L  fD  DD    CD;i 2 CD;o ln ro =ri

fD 5 1 2

CH;o ; CH;po

(5.19)

(5.20)

where DD, CD,i, CD,o, CH,o, and CH,po are diffusivity of D2 (m2/s), atomic mole concentration of D on the inside and outside surfaces of the test tube, atomic mole concentration of H on the outside surfaces of the test tube, and saturated atomic mole concentration of H, respectively.

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High Temperature Gas-cooled Reactors

The estimation of tritium permeation was carried out for the steam reforming process. It is expected that the amount of permeated tritium through the reaction tube will be reduced by counterpermeation of tritium and hydrogen. Therefore it was concluded that the amount of permeated tritium can be reduced without providing the additional countermeasures.

5.7.4 Explosion of combustible gas The explosion of combustible gas is a very severe problem to keep the reactor safe. The problem can be considered as overpressure caused by blast to the safety-related components. The simplest countermeasure is to place a long distance between the reactor and the hydrogen production system enough to mitigate the overpressure within an allowable range. However, this countermeasure brings about the rise of the construction cost of the hot gas duct and the decline of the helium gas temperature caused by heat loss from the hot gas duct, that is, the decline of the hydrogen production efficiency. Therefore, the hydrogen production system should be arranged close to the reactor to achieve economical hydrogen production. To arrange the hydrogen production system closed by the reactor, the other countermeasures, such as reducing the leak amount of combustible gas and mitigating overpressure caused by blast with barriers, should be considered. The probabilistic safety assessment (PSA) for the steam reforming process was carried out to investigate the cause of an accident on combustible gas leak and a conceptual design on a countermeasure against explosion was carried out aiming at reducing the probability of the combustible gas leak less than 1026/year. The rupture of combustible gas pipes is considered as the cause of the leak having a large impact on the reactor safety. Fig. 5.67 shows the schematic view of a coaxial pipe of combustible gas from the viewpoint of protection of the leakage. The coaxial pipe is composed of an inner pipe in which combustible gas flows and an outer pipe in which nonflammable gas such as nitrogen gas is filled. If the inner pipe is fractured, the outer pipe can protect leak of combustible gas to atmosphere and the Manhole Filled with nonflammable gas

Combustible gas

Support Outer pipe Support

Inner pipe

Figure 5.67 Schematic view of coaxial pipe for combustible gas [50].

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leak from the inner pipe can be easily detected by measuring combustible gas concentration in nonflammable gas in the outer pipe. In case of fracture of the outer pipe by detection of pressure down of nonflammable gas, hydrogen production is stopped and combustible gas in the inner pipe is purged with nonflammable gas.

5.7.5 High temperature isolation valves The isolation valve is a key component to assure safety of the HTGR hydrogen production system. It has functions to protect radioactive materials release to the hydrogen production system in case of fracture of the IHX and to prevent combustible gas ingress to the reactor in case of fracture of the chemical reactor. The HTIV used in the helium gas condition over 900 C, however, has been not fabricated for practical use yet. JAEA performed a design focusing on prevention of the valve seat deformation caused by thermal expansion, and a new coating material was developed to keep face roughness of the seat within allowable level against open and close [50,59]. An angle valve was selected from the viewpoint of workability of inner thermal insulator and the detailed structure was decided by the thermal stress analysis to prevent the valve seat deformation. Main specifications of the HTIV for the HTTRH2 are shown in Table 5.20. The coating material of the valve seat, which can keep hardness and wear resistance at high temperature over 900 C, is necessary to assure the seal performance. A new coating material was developed by adding 30 wt.%-Cr3C2 to the coating material, which is used for the valve at around 500 C. A component test was carried out with a 1/2 scale model of the HTIV, as shown in Fig. 5.68, for the HTTR-H2 to confirm the structural integrity and the seal performance of the valve seat. An experimental apparatus was composed of the 1/2 scale model, electric heaters, gas supply systems, an actuator, a concentration measurement system, and so Table 5.20 Main specifications of high temperature isolation valve for HTTR hydrogen production system [59]. Design classification

Category III valve

Fluid Temperature Pressure Flow rate Nominal size Bore Height

Helium gas 905 C 4.0 MPa 9070 kg/h 22B (O.D. 558.8 mm) 8B (I.D. 204 mm) Ab. 3 m

Material Body Seat Thermal insulator

SCPH32 Hastelloy X Isowool 1400

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High Temperature Gas-cooled Reactors

Actuator

Ground packing

Rod Cooling pipe

Disk Seat Thermal insulator

Heater

Figure 5.68 Schematic view of one-half scale model of the HTIV [50].

Actuator

F

Electric heater

P

To He gas detector

Ar gas supply system F Exhaust Cooling water Electric heater

Thermal insulator P

1/2 scale model

He gas supply system

Figure 5.69 Experimental apparatus for the HTIV [50].

on as shown in Fig. 5.69. Helium gas at 4 MPa was supplied to the 1/2 scale model, the leaked helium gas from the closed valve seat was mixed with argon gas, and the leak amount was measured by a helium gas detector. The pressure difference between supplied and leaked helium gases was 4 MPa. Before closing at 900 C, the helium gas leak rate from the valve seat at a room temperature was less than 1 cm3/s, which satisfied the design target, 4.4 cm3/s. The leak rate decreased less than 1021 cm3/s after closing at 900 C as shown in Fig. 5.70, and then it increased up to around the design target at a room temperature after opening at a room temperature. However, the leak rate became less than 1021 cm3/s when closing once again at

Leak rate of helium gas (cm3/s)

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10 4.4 cm3/s: target value

1 10-1 10-2 10-3 10-4

0

200

400

600

800

1000

Temperature of valve sheet (°C)

Figure 5.70 Leak test result of valve seat with the one-half scale model of the HTIV [50].

900 C. The following possibility is considered from the above results; the plastic deformation slightly occurred at the coating material of the valve seat by decline of the hardness when closing at 900 C, and then the face roughness of the valve seat increased when opening at a room temperature. By fitting the valve seat, the leak rate at a room temperature became less than 1 cm3/s again. The current technology can be applied to the HTTR-H2; however, the work to fit the valve seat is necessary after closing at a high temperature. Therefore the next research item is improvement of durability of the valve seat by refinement of the coating metal and so on.

5.8

Prevention technology for air ingress during a primary pipe rupture accident

5.8.1 Introduction The previous Chapter 3.5 described the basic feature of air ingress during the depressurization accident. Recently, a density-gradient-driven air ingress-stratified flow was analyzed using computational fluid dynamics (CFD) code for the Next Generation Nuclear Plant (NGNP), which is a U.S.-designed VHTR [60,61]. Furthermore, the objectives of these studies were to investigate the air ingress process and to develop a passive safe technology for the prevention of air ingress [62,63]. The experimental results show that the helium injection process is a useful method for the prevention of air ingress [6466] in the HTTR type reactor.

5.8.2 Prevention technology of air ingress in a reverse U-shaped channel 5.8.2.1 Experimental apparatus, method, and results Fig. 5.71 shows a schematic of the experimental apparatus, a cross-section of the flow channel, and a photograph of the experimental apparatus. The apparatus

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High Temperature Gas-cooled Reactors

Figure 5.71 Experimental apparatus, cross-section of a flow channel, and photograph.

Figure 5.72 Wall and gas temperature distribution for the reverse U-shaped channel.

consisted of two vertical and two horizontal circular pipes. The inner diameter of the pipe was 17 mm and the thickness of the pipe wall was 1 mm. The pipe was made of copper. The height of the vertical part was approximately 1000 mm and the length of the horizontal part was approximately 500 mm. The experimental procedure is as follows: First, the reverse U-shaped channel was filled with air. Then, one side of the vertical pipe was heated and the other side was cooled. Natural air circulation was produced by the density difference between the two vertical pipes. After a steady state was established, a small amount of helium gas was injected from the top of the channel. The velocity, temperatures of the gas and wall, and air concentration were measured during the experiment. The purpose of this experiment is to determine the possibility of controlling the natural circulation flow by injecting a small amount of helium gas. Therefore the temperature difference was set to around 40, 60, and 80 K in this experiment. Fig. 5.72 shows the wall and gas temperature distribution of the reverse Ushaped channel just after helium gas injection. Figs. 5.735.75 show the velocity

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Figure 5.73 Change in velocity of natural air circulation (29.5 mL of injected helium gas).

Figure 5.74 Change in velocity of natural air circulation (32.4 mL of injected helium gas).

change in the natural circulation of air. The injected volume of helium gas was set to 29.5, 32.4, and 35.3 mL, respectively. The temperature difference between both circular pipes was maintained at around 60 K. A small amount of helium gas was injected at 0 s. Figs. 5.765.78 show the change in air concentration for both sides of the channel. When the injected volume was 32.4 mL, the air concentration decreased rapidly. When the injected volume was 29.5 mL, helium gas moved immediately to the cooled side. However, the density difference between both vertical channels was still large. The air concentration was recovered because the natural circulation of air did not stop after the injection. As shown in Fig. 5.77, helium gas gradually diffused to the cooled side after the injection. Therefore the air concentration decreased gradually.

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High Temperature Gas-cooled Reactors

Figure 5.75 Change in velocity of natural air circulation (35.3 mL of injected helium gas).

Figure 5.76 Change in air concentration (29.5 mL of injected helium gas).

The onset time of the natural circulation of air increased with an increase in the injected volume of helium gas. The peak value of the air concentration decreased with an increase in the concentration of injected helium. The mechanism for reproducing natural circulation after helium gas injection is explained as follows: The density of gas in the heated side was usually lighter than that in the cooled side. The density of gas in the cooled side became lighter after the injection. The air and helium gas mixture moved due to molecular diffusion and very weak natural circulation. Thus the density difference increased with the elapsed time, and the natural circulation of air was reproduced. An experiment was conducted by changing the injected volume of helium gas and the temperature difference between both circular pipes. Table 5.21 and Fig. 5.79 show the relationship between the injected volume of helium gas and the onset time of the reproduction of natural circulation. The temperature difference

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Figure 5.77 Change in air concentration (32.4 mL of injected helium gas).

Figure 5.78 Change in air concentration (35.3 mL of injected helium gas).

was set to around 40, 60, and 80 K. The onset time of the reproduction of the natural circulation of air could be controlled by changing the injected volume of helium gas. The experimental results also show that the time required for molecular diffusion increased with an increase in the injected volume, and that molecular diffusion was promoted with an increase in the temperature difference. The results are described as follows: The injected volume required to stop the natural circulation of air increased with the temperature difference. The onset time of the reproduction of natural circulation increased monotonically with the injected volume. In this study, the ratio to the total volume of the channel was set in the range 3.1%12.5%.

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Table 5.21 Onset time of natural circulation flow. ΔT

V

17.7 (3.1) 20.6 (3.6) 23.6 (4.2) 26.5 (4.7) 29.5 (5.2) 32.4 (5.7) 35.4 (6.2) 38.3 (6.8) 41.3 (7.3) 44.2 (7.8) 47.1 (8.3) 50.1 (8.8) 53.0 (9.4) 56.9 (9.9) 58.9 (10.4) 61.9 (10.9) 64.8 (11.4) 67.8 (12.0) 70.7 (12.5)

40

60

80

 492 860 1470 1960 2320 2958 3470 3880 4420 4690 5250 5600 5878 6318 6940 7280 7792 8030

     810 1200 1520 2018 2320 2767 3250 3600 3920 4240 4610 4951 5140 5540

         810 997 1300 1890 2217 2670 2990 3280 3590 3790

Figure 5.79 Relationship between the volume of injected helium gas and the onset time of natural circulation.

In order to investigate the prevention method of natural circulation flow by injecting helium gas, the experiment has been performed using the apparatus which has two vertical walls. First, heavy gas is filled with the reverse U-shaped passage and the storage tank. Then, two vertical walls of the high temperature side passage are heated and cooled. Thus natural circulation flow will be generated. When the steady

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state is established, helium gas injects from the top of the passage. Natural circulation will be stopped immediately. After the time elapsed, natural circulation will be regenerated suddenly. Fig. 5.80 shows re-onset time of natural circulation. Re-onset time of natural circulation increased with increasing the amount of injecting helium gas. In order to prevent natural circulation flow, the amount of injecting helium gas increased with increasing temperature difference. Fig. 5.81 shows relationship between re-onset time of natural circulation and amount of injecting helium gas using simple reverse U-shaped tube. In this case, there is no effect on localized

Figure 5.80 Re-onset time of natural circulation in the reverse U-shaped passage.

Figure 5.81 Relationship between re-onset time of natural circulation and amount of injecting helium gas using simple reverse U-shaped tube.

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High Temperature Gas-cooled Reactors

natural convection. Therefore it takes much time to generate natural circulation. From the results obtained from these experiments, if a helium canister is placed in the upper part of the RPV or the recuperator of the GTHTR300C, air ingress from the ruptured pipe can be prevented. The amount of injecting helium gas needed to prevent the onset of natural circulation through the reactor is the same as the volume of the RPV at about 0.20.5 MPa. This is a small amount of helium gas.

5.8.2.2 Numerical analysis A numerical analysis was carried out using a 1D code. The assumptions of the numerical models were as follows: G

G

G

1D laminar flow. The diffusion coefficient of a multicomponent gas is a function of the temperature, pressure, and molar fractions of each gas component. The gas mixture obeyed the equation of state of an ideal gas.

The basic equations are derived as follows: The equation of continuity for the gas mixture: @ρ @ ðρuÞ 1 5 0: @t @x The equation of mass conservation for each gas species:   @ ðρωi Þ @ ðρωi uÞ @ @ωi 1 5 ρDi2m : @t @x @x @x

(5.21)

(5.22)

The equation of momentum conservation: ρ

  @u @u @p 1 f 1 ρu 5 2 2 ρg cosθ 2 ρujuj 1K : @t @x @x 2 De

The equation of energy conservation:       @ ρcp T @ ρucp T @ @T Lh λ 1 5 1 α ðTw 2 T Þ: @x @x Ae @t @x The equation of state for the gas: ρ p 5 RT: M

(5.23)

(5.24)

(5.25)

The diffusion coefficients Di2m for the multicomponent gas system can be obtained from the diffusion coefficients for the binary gas system and the molar fractions of each gas species, 1 2 Xi Di2m 5 Pn : (5.26) j51 Xj =Di2j The angle θ is the gradient of the flow direction, and the angle was set to 0 when the flow was upward. x is the distance from the inlet of the heated pipe. f and

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a are the frictional coefficient and heat transfer coefficient for developed laminar flow, respectively. K is the loss coefficient of the inlet and outlet of the flow. In this study, the value of K was assumed to be 1.0. The viscosity coefficient (μ) and thermal conductivity (λ) of the multicomponent gas and each component gas were obtained using Wilke’s method and Eucken’s equation. The density was calculated using the state equation [67]. An analytical model of the reverse U-shaped channel with parallel channels is shown in Fig. 5.82. It simulated the experimental conditions. The inner diameter of the pipe was 17 mm. The model consisted of two vertical pipes and one horizontal circular pipe. The height of the vertical part was approximately 1000 mm and the length of the horizontal part was approximately 500 mm. The volume of the channel was approximately 567 mL. One side of the vertical pipe was heated and its power was 80 W. The other vertical pipe was cooled by a water-cooled jacket. The temperature of the cooled pipe was maintained at the temperature of the water. The lengths of both heated and cooled parts were 500 mm. The injection point of helium gas was 1350 mm from the end of the heated pipe. The analytical area was divided into a staggered grid of 127 control volumes. The velocity was defined by the boundary surface of the control volume. The density, temperature, mass fraction, and pressure were defined at the center of the control volume. The mass conservation equation of the multicomponent gas and each component gas can be defined by the finite difference of the control volume, and the momentum conservation equation of the multicomponent gas can be defined by the finite difference of the boundary surface of the control volume. The initial condition was set with air consisting of nitrogen (79.1%) and oxygen (20.9%). The air was filled inside and outside of the reverse U-shaped channel. The

Figure 5.82 Analytical model for 1D analysis.

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High Temperature Gas-cooled Reactors

pressure in the inner and outer channels was set to atmospheric pressure. The flow velocity was set to 0 on the boundary surface of the control volume. The wall temperature of the channel was set to be equal to the experimental temperature, and the temperature distribution could be calculated by linear interpolation. The boundary condition of the wall temperature distribution of the channel is the same as that shown in Fig. 5.72. The boundary condition was set as follows: x 5 0; x 5 l XN2 5 0:791; XO2 5 0:209; XHe 5 0; p 5 p0 ; T 5 T0 ; ρ 5 ρ0 : In the 1D analysis, the temperature difference between the vertical pipes was maintained at around 57 K and a small amount of helium gas (32.4 mL; the injected volume of helium gas was about 5.7% of the total volume of the channel) was injected at 0 s. The natural circulation of air continued before injection. The change in the velocity of the natural circulation flow in the channel is shown in Fig. 5.83. The velocity decreased immediately after the injection of helium gas. After about 1800 s elapsed in the analysis (1140 s in the experiment), the velocity increased. Until natural circulation was reproduced, very weak natural circulation was observed in this analysis after the injection of helium gas. The velocity of the very weak natural circulation was 5 3 1024 to 4.7 3 1023 m/s. The time duration of molecular diffusion and the very weak natural circulation in the 1D analysis were longer than the experimental values. The velocity of ordinary natural circulation in the analysis was also larger than that in the experiment.

Figure 5.83 Velocity change with time elapsing.

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The distribution of the molar fraction of helium gas along the pipe, as determined from the analysis, is shown in Fig. 5.84. At 0 s, the molar fraction of helium gas was 1 at the injection point. Helium gas will diffuse and move due to the very weak natural circulation. The onset time of natural circulation flow was also estimated using the commercially available analysis code PHOENICS. Fig. 5.85 shows an analytical model of the reverse U-shaped circular channel. The heated and cooled water walls were provided with a constant heat flux. The boundary condition of the other wall was set to be adiabatic. The physical properties of the solid took the temperature dependence into consideration. The physical properties of the fluid took the temperature and pressure dependence into consideration. The number of calculation cells was 86 3 14 3 78 5 0.85 million. Although PHOENICS can use the k-ε (KECHEN) turbulence model, we did not use the model in the analysis because the analysis condition was in the laminar flow region. In the 3D analysis, the temperature difference between the vertical pipes was maintained at around 57 K and a small amount of helium gas (32.4 mL; the injected volume of helium gas was about 5.7% of the total volume of the channel) was injected at 0 s. From Fig. 5.72, the average temperature was 74 C for the heated side channel and 17 C for the cooled side channel. Then, the boundary condition was maintained at 74 C for the heated part and 17 C for the cooled part. The other parts of the reverse U-shaped channel, such as the horizontal part and the inlet and outlet parts of the channel, were set to an adiabatic boundary condition. The natural

Figure 5.84 Distribution of the molar fraction of helium gas during the experiment.

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High Temperature Gas-cooled Reactors

Figure 5.85 Analytical model for 3D analysis.

circulation of air continued before injection. The velocity change in the natural circulation flow in the channel is shown in Fig. 5.86. The velocity decreased immediately after helium gas injection. After 1230 s elapsed in the analysis (1140 s in the experiment), the velocity increased. Until the reproduction of natural circulation, very weak natural circulation was also observed in this analysis after the injection of helium gas. The velocity of the very weak natural circulation was on the order of 1 3 1024 to 1 3 1023 m/s. The duration time of molecular diffusion and the very weak natural circulation in the 3D analysis are in good agreement with the experimental values. The velocity of ordinary natural circulation in the analysis is also in good agreement with the experimental values. The distribution of the molar fraction of helium gas along the pipe, as determined using the 3D analysis, is shown in Fig. 5.87. Helium gas diffuses and moves due to the very weak natural circulation. Fig. 5.88 shows a contour line of the molar fraction of helium gas. Helium gas moved from the injection point to the cooled side of the reverse U-shaped pipe.

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Figure 5.86 Velocity change with time elapsing.

Figure 5.87 Distribution of molar fraction of helium gas during the experiment.

Fig. 5.89 shows the SCAD system for the GTHTR300C system [68] proposed by JAEA. If a helium canister is placed in the upper part of the RPV or the recuperator, air ingress from the ruptured pipe can be prevented. The amount of helium gas needed to prevent the onset of natural circulation through the reactor is the same as the volume of the RPV at 0.20.5 MPa. This is not a large amount of helium gas.

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High Temperature Gas-cooled Reactors

Figure 5.88 Contour lines of the molar fraction of helium gas.

Figure 5.89 SCAD system for GTMHR300 proposed by JAEA. [68]

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5.8.2.3 Onset time of natural circulation flow through the apparatus Authors have reported on the mixing process of two component gases through natural convection and molecular diffusion in a stable stratified fluid layer [69]. According to the report, the mixing process through molecular diffusion in the vertical stratified fluid layer was significantly affected by localized natural convection induced by the slight temperature difference between both vertical walls. Authors have also reported that transport phenomena due to molecular diffusion were affected by not only localized natural convection but also by natural circulation of the gas mixture. Localized natural convection may affect an onset time of natural circulation. In order to predict or analyze the air ingress phenomena during a depressurization accident in a HTGR, it is important to examine the influence of localized natural convection and molecular diffusion on the mixing process. In general, mixing processes of two component gases in a vertical stable stratified fluid layer are often governed by molecular diffusion. When a stable stratification is formed in a vertical slot with two component gases, which have different densities, the rate of transportation will be different, as determined by a mutual diffusion coefficient. On the other hand, it is expected that natural convection will occur in the vertical slot when one sidewall is heated and the other sidewall is cooled. When a stable stratification is formed with the two component gases and the two vertical parallel walls of the slot are kept at different temperatures, the transport process of the gases becomes more complex. In this case, the heavier gas diffuses into the lighter gas. In addition to that, these gases will also be transported by localized natural convection. Both phenomena may occur at the same time during the air ingress process of the primary pipe rupture accident. According to previous experiments [70,71], molecular diffusion and natural convection will occur simultaneously in the annular passage between the inner barrel and the watercooled jacket. The range of the Rayleigh number based on the width of the annular passage is about 0 , Rad , 3.26 3 105 and Rad , 1.56 3 106, respectively. The Rayleigh number based on the width of the annular passage of the HTTR or the GTHTR300C will be larger than two digits of Rayleigh number of the simulated apparatus. Therefore it is necessary to know which phenomenon is dominant in the mixing process of two component gases in a vertical stable stratified fluid layer. It is also important to quantitatively evaluate the influence of natural convection on the mixing processes due to molecular diffusion. In this chapter, the onset time of natural circulation of air was discussed in terms of the results obtained from the previous experiments and the present experiment.

5.8.2.4 Onset time of natural circulation Fig. 5.90 shows the onset time of natural circulation against the wall temperature of the high temperature side passage. Fig. 5.91 shows the onset time of natural circulation obtained by the experimental results which was obtained by using three apparatus. Three apparatus are the reverse U-shaped tube [64,70,71], three parallel channels [72], and vertical parallel walls [65,69]. The height of the heated part is less than 1 m. As shown in Fig. 5.91, the onset time of natural circulation becomes long when the phenomenon is governed mainly by molecular diffusion. When not only the

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High Temperature Gas-cooled Reactors

Figure 5.90 Onset time of natural circulation against wall temperature of the hightemperature side passage.

Figure 5.91 Onset time of natural circulation against wall temperature of the hightemperature side of the reverse U-shaped passages.

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localized natural convection but also natural circulation was generated, the onset time of natural circulation becomes short. Therefore, it is important to know the position of generating of the localized natural convection and the localized natural circulation.

5.8.3 Basic feature of air ingress phenomena during a horizontal pipe break accident 5.8.3.1 Introduction The other experiment for a horizontal pipe break case is planned using an apparatus, which is shown in Fig. 5.92. Air ingress scenario in the case of the horizontal pipe break of the GTHTR300C is as follows: After the pipe ruptures, air will flow into the bottom part of the RPV by the counter-current flow. The density stratified fluid layer will be formed. Buoyancy will produce between the hot and cold legs. As the buoyancy will be small, the natural circulation flow will not produce under the condition of this density distribution. Thus air will be transported to the reactor core by mainly molecular diffusion. However, from the results obtained in these experiments, air will be transported to the reactor core by localized natural convection. In the configuration of the HTTR, a vertical channel existed between the reactor core and the pipe rupture part. So, it took much time to generate natural circulation of air. In the configuration of the GTHTR300C, the vertical path does not exist between the reactor core and the pipe rupture part. Therefore air may be transported to the reactor core earlier. The onset time of the natural circulation of air and the amount of infiltration air during the accident will be greatly affected to the generated position and strength of the localized natural convection in the pressure vessel. If localized natural convection occurs inside the channel, it will be difficult to estimate not only the density change of gas mixture but also the onset time of natural circulation through the reactor. Anyway, after the time elapses, natural circulation may occur suddenly. In order to research the mixing process of

Figure 5.92 Apparatus simulated horizontal pipe rupture accident in GTHTR300C.

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High Temperature Gas-cooled Reactors

two-component gas when the horizontal primary pipe of the GTHTR300C is ruptured, experiment and numerical analysis are planned.

5.8.3.2 Experimental apparatus Fig. 5.93 shows a schematic drawing of an experimental apparatus. The experimental apparatus consisted of a double coaxial cylinder and a horizontal double coaxial pipe. The material of these parts was stainless steel. The outer pipe of the horizontal double coaxial pipe was connected to the outer cylinder of the double coaxial cylinder, and the inner pipe of the horizontal double coaxial pipe was connected to the inner cylinder of the double coaxial cylinder. The outer pipe of the horizontal double coaxial pipe was 34 mm in outer diameter and 27.2 mm in inner diameter. The inner pipe was 21.7 mm in outer diameter and 16.1 mm in inner diameter. The outer cylinder of the double coaxial cylinder was 410 mm in height and 283.4 mm in diameter. The inner cylinder was 255 mm in height and 139.8 mm in diameter. Four cartridge heaters were installed in the inner cylinder, the cartridge heater was 300 mm in length, 12.8 mm in outer diameter, and 170 mm in effective heating portion, with a rated voltage of 100 V and rated capacity of 100 W. There was a helical waterway in the outer cylinder, and the outer cylinder was cooled by flowing water. To prevent heat radiation, insulation (ceramic type) with a height of 200 mm and a thickness of 23.5 mm was installed on the inner cylinder. Therefore the gap between the inner cylinder and the outer cylinder was 46.7 mm. For measurement of the gas concentration, an ultrasonic gas concentration meter was used. This meter utilized the property that the sound velocity changes as the gas concentration changes. The ultrasonic wave was sent to the sampled gas, and the concentration of the two-component gas could be measured from the time until it reaches the receiving sensor. The gas concentration measurement and the

Figure 5.93 Experimental apparatus.

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thermocouple installation positions are shown in Ref. [73]. The gas was sampled with a microtube pump from the sampling ports provided in the bottom part, upper part, and horizontal double coaxial pipe of the experimental apparatus, and the concentration was measured. K-type thermocouples were used for temperature measurement. The temperature measurement error accuracy of this thermocouple was 6 1.5 K. The thermocouple was inserted through a pore of 1.2 mm in diameter of a compression fitting provided on the outer cylinder. Six thermocouples were installed at the 5-mm position from the outer cylinder. Six thermocouples were installed at the 16.5-mm position from the outer cylinder. Six thermocouples were installed at the 28-mm position from the outer cylinder. A total of 18 thermocouples were installed, and gas temperature changes in the outer channel were measured. To measure the gas temperature at the bottom of the experimental apparatus, six thermocouples were installed at 25 mm from the bottom of the apparatus. To measure the gas temperature at the top of the experimental apparatus, eight thermocouples were installed on the ceiling of the experimental apparatus. Two thermocouples for cooling-water temperature measurement were installed at the water-cooling jacket entrance. The heater temperature was measured with a built-in thermocouple. A constant-temperature-type thermal anemometer was used to measure the flow velocity. The flow velocity was measured by attaching an anemometer to the valve installed on the horizontal double coaxial tube.

5.8.3.3 Experimental method The experimental procedure was as follows: First, the experimental apparatus was filled with helium gas. The inner cylinder was heated, and the outer cylinder was cooled by flowing water. Heat inputs were 36 W (1.22 kW/m2 ), 144 W (4.89 kW/m2 ), and 324 W (11.0 kW/m2 ), and the water flow rate was constant. After the heater temperature was reached at steady state, the valves of the inner pipe and outer pipe were opened at the same time to simulate a primary pipe rupture accident. During the experiment, the temperature and mole fraction of the two-component gases were measured. The flow velocity at the outlet of the horizontal double coaxial pipe was also measured.

5.8.3.4 Experimental results Fig. 5.94 shows gas temperature changes at various places under the heat input conditions of 324 W. Fig. 5.95 shows heater temperature under each condition. Fig. 5.96 shows the inlet and outlet cooling-water temperature. The abscissa shows the elapsed time, and the ordinate shows temperature. The opening time of the valves was 0 min. Comparing the three conditions, the gas temperature near the inner cylinder rose greatly as the heat input increased, but the gas near the lower part of the outer cylinder did not rise too much. Therefore as the heat input increased, the temperature difference of the air inside the apparatus increased, and it seems that natural convection through the apparatus is likely to occur. Further, as

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Figure 5.94 Gas temperature changes (heat input 324 W).

Figure 5.95 Heater temperature changes.

shown in Fig. 5.94, the gas temperature at the inner cylinder fluctuated up to 6 2.0 K during the experiment. This small temperature fluctuation indicates the occurrence of local natural convection, which is thought to promote the mixing of the gas. Slight temperature fluctuation of the gas was confirmed also on the outer cylinder. The temperature fluctuation of the gas near the outer cylinder increased when going to the upper part, and the strength of the local natural convection is considered to increase accordingly.

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Figure 5.96 Cooling-water temperature (heat input 324 W).

Figure 5.97 Air mole fraction changes (heat input 324 W).

Fig. 5.97 shows the air mole fraction at various places under the heat input condition of 324 W. The horizontal axis shows the elapsed time, and the vertical axis shows the air mole fraction. The opening time of the valves was 0 min. The range where the air mole fraction was 0.5 or less is a prediction because it is impossible to measure because of the characteristics of the measuring instrument. As shown in Fig. 5.97, after opening the valves, the air mole fraction at the inner pipe rose rapidly compared with the other points. It is predicted that after opening the valves, the air enters mainly from the inner pipe. Furthermore, comparing results shows

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that the higher the heat input, the shorter the time until the inside of the experimental apparatus is completely filled with air. Therefore the temperature inside the inner cylinder affects the amount of air entering. Fig. 5.98 shows the flow velocity at the outer pipe of the horizontal double coaxial pipe under the heat input conditions of 324, 144, and 36 W. The abscissa shows the elapsed time and the ordinate shows the flow velocity. The opening time of the valves was 0 min. In the result for a heat input of 324 W, immediately after opening the valves, the flow velocity at a horizontal double coaxial pipe increased rapidly because air entered into the experimental apparatus by the countercurrent flow. Subsequently, the flow velocity decreased. It is considered that the stable density stratified fluid layer was formed, and the flow velocity decreased. Afterward, the flow velocity increased again. As time elapsed, the gas mixture in the inner cylinder was heated, and the gas mixture near the outer cylinder was cooled. Thereby, a natural circulation flow that circulated through the apparatus occurred. Therefore the flow velocity should increase. After this, the flow velocity decreased again and reached a constant value. Because of the characteristics of the thermal anemometer used, air was detected to be lower in flow velocity than helium gas. Thus this tendency was seen when the experimental apparatus was filled with air. From this result, a stable density stratified fluid layer of twocomponent gas should be formed when the primary pipe breaks. However, the natural circulation flow that circulates from the reactor core through the RPV will generate afterward. Further, at the heat inputs 144 and 36 W, there was no tendency similar to that at the heat input of 324 W. The natural circulation flow that circulates through the apparatus did not occur when the heat input was low, and the entry of air was also delayed.

Figure 5.98 Flow velocity changes at the outer pipe of the horizontal double coaxial pipe.

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Experiments were conducted to investigate the air ingress process by using an experimental apparatus in which a horizontal double coaxial pipe was connected to a double coaxial cylinder when a primary pipe of the GTHTR300C was broken. 1. The temperature fluctuation of the gas near the outer cylinder increased when going to the upper part, and the strength of the local natural convection increased accordingly. 2. After opening the valves, the air mole fraction at the inner pipe rose rapidly compared with the other points. Therefore the air entered mainly from the inner pipe. 3. The higher the heat input, the shorter the time until the inside of the experimental apparatus was completely filled with air. Therefore the temperature inside the inner cylinder affects the amount of air entering.

A stable density stratified fluid layer of two-component gas forms when the primary pipe breaks. However, the natural circulation flow that circulates from the reactor core through the RPV will generate afterward. By clarifying the relationship between the reactor core temperature and the onset time of natural circulation, it can be determined how quickly to prevent graphite oxidation at the depressurization accident.

5.9

Advanced fuel technology for high burnup

5.9.1 Introduction The inherently safe HTGR is focused on a supersafe nuclear system after the FukushimaDaiichi Nuclear Power Plant Accident. JAEA has been progressing to design HTGR fuels for not only small-type practical HTGRs but also VHTR proposed in GIF, which can be utilized for various purposes with high temperature heat at 750 C950 C. To increase economy of these HTGRs, JAEA has been upgrading the design method for the mass-produced HTGR fuel, which can maintain their integrities at the burnup of three to four times higher than that of the conventional HTTR fuel, so-called High Burnup. In addition, a concept of a plutonium (Pu)-burner HTGR, so-called Clean Burn, is being developed which targets a burnup of 10 times or more than that of ordinary uranium fuel of 500 GWd/t and consumes surplus Pu inventory in Japan with high proliferation resistance by employing an IMF and a tightly coupled fuel reprocessing and fabrication plants. Furthermore, a waste loading method for direct disposal of the spent HTGR fuel is investigated, which can achieve “Reduction of High-Level Radioactive Waste” and footprint in a geological repository compared with LWR because of higher burnup, higher thermal efficiency, and less trans uranium (TRU) generation of HTGR. As for the advanced fuel system, the following three items have been researched: 1. High burnup, 2. Plutonium burner, and 3. Reduction of high-level radioactive waste.

JAEA has been progressing to design HTGR fuels not only for small-type practical HTGRs but also for VHTR proposed in GIF, which can be utilized for various purposes with high temperature heat at 750 C950 C. To increase economy of

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these HTGRs, JAEA has been upgrading the design method for the HTGR fuel, which can maintain their integrities at the burnup of three to four times higher than that of the conventional HTTR fuel. Specification of the fuel for each HTGR is summarized in Table 5.22 [74]. The design details of the HTR50S and the GTHTR300 are described in Section 5.1. The HTTR employs so-called pin-in-block-type fuel element consisting of fuel rods and hexagonal graphite block. TRISO (Tristructural isotropic)-coated fuel particle (CFP) with 12 kinds of low enriched uranium dioxide (UO2, 3.4%9.9% of 235 U; 6% in average) fuel kernel is employed to attain the burnup of 22 GWd/t in average and 33 GWd/t in maximum. The HTR50S has a concept to attain higher burnup of 100 GWd/t (in maximum) than that of HTTR, which employs the so-called extended burnup (EBU) TRISOCFP, consisting of a smaller UO2 kernel, thicker buffer, and silicon carbide (SiC) coating layers than those of HTTR type one to reduce probability of CFP failure due to not only kernel migration but also internal gas pressure as described below. The HTR50S also employs pin-in-block type fuel element. The HPC HTGR is also a small-sized HTGR, which will be adopted same TRISOCFP to be used by 100 GWd/t in maximum as that for the HTR50S [75]. In addition, as a characteristic feature of the HPC HTGR, the monolithic fuel rod made with the SiC matrix is adopted in order to enhance an extreme safety against in-core oxidation accident. The GTHTR300 also adopts the high burnup TRISO-CFP targeting at 140 GWd/t in average and 155 GWd/t in maximum. One of the characteristic points of this fuel is larger diameter of TRISO-CFP than any other HTGR in the world by adopting thicker buffer layer to prevent failure due to the internal gas pressure and kernel migration. The Pu-burner HTGR is proposed to reduce plutonium inventory by combustion using the inherently safe HTGR as focused on a supersafe nuclear system after the Fukushima-Daiichi Nuclear Power Plant Accident. In order to burn plutonium efficiently and in large quantities, the Pu-burner HTGR targets a burnup of 10 times or more than that of the ordinary uranium fuel of 500 GWd/t [76]. In addition, the fabricated TRISO-CFP also had been confirmed to be high-reliability, that is, 1/100 times smaller damage ratio over the conventional one, through the HTTR operation [77]. Therefore so-called 3S-TRISO, that is, TRISO-CFP with safety, security, and safeguards, is developed for the Pu-burner HTGR. This fuel employs yttriastabilized zirconia (YSZ), which is chemically inactive from the viewpoint of not only safety at the time of direct disposal but also nuclear nonproliferation [78], as an inert matrix of plutonium dioxide (PuO2) fuel kernel, thereby strengthening nuclear proliferation resistance by inert fuelization. Furthermore, zirconium carbide (ZrC) is combined to suppress the rise in internal pressure of carbon monoxide (CO) gas, which is the main cause of internal pressure failure during irradiation.

5.9.2 Design of high burnup fuel Retentiveness of fission products (FPs) within TRISO-CFP is important in the fuel safety design to keep their releases to the primary coolant below an

Table 5.22 Specifications of HTGR fuels development by JAEA [74]. Reactor

HTTR

HTR50S

HPC HTGR

GTHTR300

Pu-burner HTGR

30 MW 850 C 950 C/395 C 33 GWd/t 660 days 2.5 MW/m3

50 MW 750 C/325 C 100 GWd/t 730 days 3.5 MW/m3

165 MW 750 C/325 C 100 GWd/t 1200 days 3.8 MW/m3

600 MW 850 C/587 C 155 GWd/t 730 days 5.4 MW/m3

600 MW 850 C/587 C 625 GWd/t 430 days 56 MW/m3

UO2 3.4%9.9% 235U (12 kinds) 600 μm No Ld PyC/60 μm Hd PyC/30 μm

UO2 4.3%9.4% 235U (3 kinds) 500 μm No Ld PyC/95 μm Hd PyC/40 μm

UO2 ,20% 235U

UO2 14% 235U

500 μm No Ld PyC/95 μm Hd PyC/40 μm

550 μm No Ld PyC/140 μm Hd PyC/25 μm

PuO2-YSZ 62% 239Pu in Pu composition 400 μm ZrC/10 μm Ld PyC/95 μm Hd PyC/40 μm

SiC/25 μm Hd PyC/45 μm 920 μm

SiC/35 μm Hd PyC/40 μm 920 μm

SiC/35 μm Hd PyC/40 μm 920 μm

SiC/40 μm Hd PyC/25 μm 1010 μm

SiC/35 μm Hd PyC/40 μm 840 μm

A3-3a Cylindrical with graphite sleeve OD26 3 ID10 3 H39 mm

A3-3 Cylindrical with graphite sleeve OD26 3 ID10 3 H39 mm

SiC Monolithic

A3-3 Monolithic

A3-3 or SiC Monolithic

OD26 3 ID10 3 H83 mm

30%

30%

60%

OD24 3 ID9 3 H83 mm with sheath 1 mm thick 29%

OD24 3 ID9 3 H83 mm with sheath 1 mm thick 30%

Items Thermal power Outlet/inlet Temp. Max. burnup/ periods Average power density

Coated fuel particle Fuel kernel Enriched fissile metal Diameter Oxygen getter Buffer coating/thick TRISO coating/ thick

CFP diameter

Fuel compact Matrix material Shape Dimension Particle packing fraction a

A3-3 means the mixture of 64 wt.% of natural graphite, 16 wt.% of petroleum graphite, and 20 wt.% of phenol resin (the value means weight percent before carbonization).

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acceptable level. From this viewpoint, JAEA settles the basic design criteria for the fuel as the following [79]: G

G

G

G

G

As-fabricated fuel failure fraction shall be minimized. Significant additional failure shall be avoided during the operation. The following meet this criterion: Fuel shall not fail systematically under normal operating condition. Fuel temperature shall be limited below 1600 C under anticipated occasional occurrences (AOOs). Fuel burnup is limited in the basis of the results of the irradiation test.

Since the HTTR is the first HTGR constructed in Japan, a RPV and major primary cooling system such as primary pressurized water cooler and intermediate heat exchanger are located inside a containment vessel. Therefore as-fabricated fuel failure fraction is limited lower than 0.2% from the viewpoint of limit of off-site exposure during the normal operation [79]. Furthermore, fuel burnup is limited to 33 GWd/t in maximum in the basis of the results of irradiation tests. For the GTHTR300, the HTR50S, and the HPC HTGR, a confinement structure will be adopted to the reactor building in spite of the containment vessel applied for the HTTR, in order to achieve economy of the reactor construction. Therefore in case of the GTHTR300 fuel, as-fabricated fuel failure fraction is limited below 5 3 1024 based on experiences of the HTTR first-loading fuel fabrication. Asfabricated fuel failure fractions for the HTR50S and the HPC HTGR are not be published yet, however, would be limited as same level as the GTHTR300 fuel. For the safety design of the HTTR fuel, two systematic failure modes are taken into consideration [79]: kernel migration [80] and corrosion of SiC coating by palladium [81]. On the other hand, it is known that UO2-TRISO CFP shall be failed by increasing internal pressure of TRISO due to CO gas and fission gases generated by fission. Under high burnup over 100 GWd/t, internal pressure failure becomes dominant, although this failure mode was not applied for the HTTR fuel because of its lower burnup, 33 GWd/t in maximum. Therefore in the safety design of the HTR50S, HPC HTGR, and GTHTR300’s TRISO-CFPs, the following tendencies are seen: G

G

Smaller diameter of UO2 fuel kernel and thicker buffer layer to reduce internal gas pressure. Thicker SiC coating layer to increase integrity of TRISO as a pressure vessel.

5.9.3 Upgrade technologies for high burnup 5.9.3.1 Fuel design The reference specification of the upgraded TRISO-CFP, in case of EBU TRISOCFP, has been determined according to the following requirements [82]: G

The diameter of the new TRISO CFP was determined to be same as the HTTR, because of minimizing fuel failure fraction in the HTTR fuel compaction condition and keeping accuracies in fuel compact inspections.

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G

G

G

G

425

The diameter of UO2 kernel was determined to be smaller than that of the HTTR in order to increase the volume of buffer layer for avoiding the pressure vessel failure of the CFP. A reference diameter of 500 μm for EBU TRISOCFP was settled considering the fuel kernel fabrication process and a past experience on 550 μm kernel fabrication [83]. The thickness of buffer layer was determined to be larger than that of the HTTR in order to avoid two modes of failures due to amoeba effect [80] and to the internal gas pressure. The thickness of SiC layer shall be increased to enhance the integrity of TRISO CFP as the pressure vessel. It was determined to be larger than that of the HTTR considering the fuel kernel fabrication process and a past experience on 35 μm thick fabrication [83]. Each IPyC and OPyC layer is coated with similar thickness to the HTTR to maintain isotropic quality by the fabrication.

To evaluate failure probability of the above TRISO-CFP due to internal gas pressure under the irradiation, JAEA has developed the calculation codes such as FIGHT [84] and Code-B-2 [85]. The essential features of both fuel failure models are based on the following [84]: G

G

G

G

The model calculates failure probability of each coating layer. The failure probability of each coating layer is described by the Weibull distribution with microscopic surface flaws considered to be critical failure initiation sites. The failure criterion of each coating layer is determined based on experimental observations accumulated from irradiation tests. The stresses acting on the coating layers of the fuel particle are assumed to be caused by pressure of fission gases and CO gas from UO2 kernel and by fast neutron-induced shrinkage of PyC layers (CO generates by excess oxygen liberated during fissioning of UO2 reacting with carbonaceous buffer coating.). The maximum stress is evaluated by the rigid SiC model with spherical shell because recently fabricated fuel particles have good sphericity. Two types of failed particles are categorized, that is, the through-coatings failed particle and the SiC-failed particle. The through-coatings failed particle results from IPyC, SiC, and OPyC layer failure (i.e., an exposed fuel kernel). The SiC-failed particle has failed IPyC and SiC layers but an intact OPyC layer. The gaseous FPs are released from through-coatings failed particle. On the other hand, since OPyC layer is capable of retaining the gaseous FPs, the SiC-failed particle does not release the gaseous FPs. The failure probability that an as-fabricated SiC-failed particle becomes a through-coatings failed particle is also modeled.

The fracture strength distribution of coating layers is expressed by a Weibull distribution. The failure probability of each layer is calculated by the following basic equation: 

   σi ðtÞ mi fi ðtÞ 5 1 2 exp 2ln2 3 ; σ0;i

(5.27)

where fi(t), σi(t), σ0,i, and mi are the failure probability of the i-layer at irradiation time t, the stress on the i-layer at irradiation time t, the strength of the i-layer, and the Weibull modulus for the i-layer strength, respectively.

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The failure probability of the coating layer is modeled as follows: G

G

SiC layer fails by internal gas pressure in high burnup irradiation condition. The compressive stress caused by IPyC and OPyC layers shrinkage mitigates the stress on SiC layer. When the OPyC layer is intact, the failure probability of SiC layer (fSiC) is evaluated based on the tensile stress, which is a balance of the stress due to internal gas pressure and the compressive stresses by both IPyC and OPyC shrinkage. OPyC layer fails only by tensile stress. When SiC layer, which is inside of OPyC layer, is intact, the failure of OPyC layer does not occur (f 0OPyC 5 0) because it is supported by the stronger SiC layer. In the SiC-failed particle, the failure of OPyC layer occurs by the internal gas pressure (f 0OPyC ).

Based on above-mentioned failure behavior, probability that the intact particle becomes the through-coatings failed particle (FTG) can be expressed as the following: FTC 5 f 0IPyC 3 fSiC 3 f 0OPyC ;

(5.28)

where f 0IPyC and f 0OPyC are the failure probabilities of IPyC and OPyC layers, respectively. Fig. 5.99 shows the relationship between dimensions of the buffer layer, SiC layer, and the fuel kernel of each TRISO-CFP and the probability to lose its intactness against the pressure vessel failure at the burnup of 120 GWd/t for instance. In this calculation, the failure probability of CFP was settled at 1 3 1026, and other calculation conditions were as follows: 1300 C of irradiation temperature, 1460 effective full power days (EFPDs), and 3.3 3 1025 m22 of the fast neutron

Figure 5.99 Relationships between kernel diameter, buffer layer thick, and 1 3 1026 of failure probability (φfail) at 1300 C and 120 GWd/t for HTTR, EBU, and GTHTR300 types TRISO-CFPs [74].

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fluence (E . 0.29 fJ). Regardless of original specification of every type of CFP, it was assumed that the corresponding amount of 235U was loaded and burned by fission to 120 GWd/t in order to make the same internal gas generating condition. Each dotted line in the figure shows a boundary of probability at 1 3 1026 for each SiC layer thick. As the result of the fabrication, EBU-type TRISO-CFP keeps intactness with the failure probability at 1 3 1026, and GTHTR300 type would not fail in this irradiation condition. The HTTR type would fail with higher probability than others in this irradiation condition since it was designed for use with the burnup up to 33 GWd/t. The graphite sleeve of the HTTR fuel works to prevent oxidation of the fuel compact from impurities in the primary coolant under normal operation, and also from oxygen and/or humidity under the air/water-ingress accident. On the other hand, the gap distance between outer surface of fuel compact and inner surface of the sleeve increases during the irradiation because the shrinkage rate of fuel compact matrix by fast neutron is larger than that of graphite [86]. Then, the difference of temperature due to this gap would become around 90 C in case of the HTTR [87], which could cause to decrease the fuel temperature as the design limit. Therefore HPC HTGR and GTHTR300 fuel employ the monolithic fuel rod, which has a thin scale made of same material as fuel compact matrix on the surface of fuel compact in spite of the sleeve to enhance the heat removal and to decrease fuel temperature.

5.9.3.2 Irradiation test The EBU TRISO-CFP has been fabricated based on the HTTR fuel technology and irradiated up to 100 GWd/t in target. As the result of the fabrication, dimensions of newly designed CFP including the tolerance were within the failure probability of 1 3 1026 as shown in Fig. 5.100, and it was suggested that the new CFP could fabricate successfully to maintain its intactness at 100 GWd/t even if considering the fabrication tolerance. The irradiation test with the high burnup fuel has performed in the irradiation conditions at 1.0 3 1018 m22/s of thermal neutron flux and 1050 6 100 C of irradiation temperature corresponding to the HTGR normal operating condition. Finally, it took for 403 EFPDs of irradiation duration and the calculated level of burnup has comprised about 93.3 GWd/t to the end of irradiation. Release rate to birth rate ratio

Figure 5.100 Cross-sections of HTTR and EBU TRISO-CFPs [74].

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High Temperature Gas-cooled Reactors

(R/B) of 88Kr indicated that two TRISO CFPs in three fuel compacts were failed additionally during the irradiation, which was less than the number of as-fabricated fuel failure, that is, three failed CFPs in three fuel compacts. It is known that the dimensional shrinkage of a fuel compact is varied depending on the fast neutron fluence (E . 29 fJ), which is almost linearly increased at the beginning of irradiation (B1.5 3 1025 m22) and gradually decreased after that up to 3 3 1025 m22 following the approximate equation described typically as the secondary expression. Also, an experimental datum indicated that almost no influence on the fuel compact shrinkage was observed in the temperature range from 800 C to 1700 C [13]. It is important to extend irradiation data of fuel compact to the condition for the practical use over 3 3 1025 m22 (E . 29 fJ). For compaction, two kinds of fuel compact matrix were used: one was the conventional HTTR’s A3-3 matrix; another consisted of newly chosen natural- and petroleum-graphite powders supplied by Japanese graphite fabricators and the conventional binder, of which the weight ratio was as same as the HTTR. Each new graphite powder was selected according to the properties of conventional powders used for the HTTR. As a nondestructive test with irradiated fuel compact, irradiation shrinkage rate as a function of fast neutron fluence was evaluated as shown in Fig. 5.101. Each dimension of fuel compact before and after irradiation was measured by micrometer. As a comparison, the figure shows the data of the 94F-9A capsule irradiated with the HTTR first-loading fuel up to 60 GWd/t and of the 91F-1A capsule with high burnup-type HTTR TRISO-CFP up to 90 GWd/t in Japan Material Testing Reactor (JMTR). As the result, shrinkage rate of fuel compact using conventional matrix material shows good agreement with data obtained in the HTTR project. Also, fuel compacts using new graphite matrix perform good tendency of shrinkage by irradiation. Finally, an equation of shrinkage rate of

Figure 5.101 Irradiation shrinkage rate of fuel compact as function of fast neutron fluence [74].

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fuel compact matrix up to 5.6 3 1025 m22 of fast neutron fluence (E . 29 fJ) was determined as the following; ΔD ð%Þ 5 0:0722φ2 2 0:798φ; D

(5.29)

where φ is the fast neutron fluence. These data will contribute to the future thermal-hydraulics designs not only for HTR50S adopting the pin-in-block type fuel element (with sleeves) but also for GTHTR300 with the monolithic-type graphite-matrixed fuel element.

5.9.4 Future study plan The high burnup fuel consists of the upgraded TRISO-CFP for high burnup use, and the fuel compacts with two kinds of graphite matrices, the conventional and the newly developed ones. For upgrading the design method of TRISO-CFP for enhancing economy of the HTGR, its failure probability was evaluated by the socalled the internal gas pressure failure model. To demonstrate the integrity of the EBU TRISO-CFP and irradiation performance of fuel compact matrix, irradiation tests and PIEs were conducted. As the latest result, the EBU TRISO-CFP has successfully demonstrated its integrity at the burnup nearly 100 GWd/t, which is three to four times higher than that of the HTTR fuel (33 GWd/t). In addition, through the PIEs, the extended data of shrinkage rates of A3-3 and newly selected graphite matrices for the fuel compact as a function of fast neutron fluence over 5 3 1025 m22 (E . 0.29 fJ) were obtained successfully, which will contribute to the future thermal-hydraulics designs of the HTGR core using graphite matrix. Furthermore, as the future R&Ds needed for the high burnup HTGR fuel, adoption of larger fuel kernel could be essential to extend lifetime of the fuel instead of increasing not only fissile enrichment but also particle packing fraction of fuel compact. As one of candidates for the advanced high burnup fuel, ZrC-coated UO2 TRISO-CFP would have a strong potential to extend burnup. Other advanced TRISO-CFP concepts under development include the SiC-matrix fuel to essentially eliminate oxidation related failure mechanisms during severe accidents and provide even higher levels of defense-in-depth and inherent safety.

5.10

Advanced fuel for plutonium burner

5.10.1 Introduction About 44 t of recovered plutonium had accumulated in Japan by 2012 [88]. It will further increase by about 8 t per year after the operation of a reprocessing plant at Rokkasho gets restarted. However, there is not enough capability to consume plutonium by Japan’s nuclear reactors yet. An innovative plutonium burner concept based on HTGR technology, “Clean Burn,” is proposed by JAEA [76]. That is expected to be as an effective and safe

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method to consume surplus plutonium accumulated in Japan. A similar concept proposed by General Atomics (GA), Deep Burn [89], cannot be introduced to Japan because of its adopting highly enriched plutonium, which shall infringe on a Japanese nuclear nonproliferation policy according to JapanUS reprocessing negotiation. The Clean Burn concept can avoid this problem by employing an IMF and a tightly coupled fuel reprocessing and fabrication plants. Both features make it impossible to extract plutonium alone out of the fabrication process and its outcomes. As a result, the Clean Burn can use surplus plutonium as a fuel without mixing it with uranium matrix. Thus surplus plutonium alone will be incinerated effectively, while generation of plutonium from the uranium matrix is avoided. High neutronic performance, that is, achievement of burnup of about 500 GWd/t and consumption ratio of 239Pu reaching to about 95%, is also assessed. Furthermore, reactivity defect caused by the inert matrix is found to be negligible. It is concluded that the Clean Burn concept is a useful option to incinerate plutonium with high proliferation resistance. The features of Clean Burn are summarized as follows: G

G

High proliferation resistance by employing IMF and a tightly coupled plant of fuel reprocessing and fabrication, and High plutonium incineration ability.

The main objective is to incinerate fissile plutonium “Cleanly.” For fissile nuclides, 241Pu has the short half-life of 14.4 years. On the contrary, the half-life of 239 Pu is a long term of 24,100 years. For the risk of proliferation after geological disposition, only 239Pu is a problem. Thus the target of the Clean Burn concept is to incinerate around 95% of loaded 239Pu.

5.10.2 Fuel fabrication process of Clean Burn The plutonium must be mixed with the recovery uranium, and the enrichment must be less than 50 wt.% in the present situation in the Japanese nuclear fuel cycle environment due to the JapanUS reprocessing negotiation. Thus the plutonium consumption efficiency is defected significantly by the plutonium generation from the fuel matrix of uranium. If an alternative reprocessing system to treat pure plutonium without uranium mixing would be proposed, the system should have the higher proliferation resistance than the existing one to be accepted in Japanese nuclear fuel cycle environment. In this context, an innovative plant system [90] combining the LWR spent fuel reprocessing and HTGR fuel fabrication with high proliferation resistance is proposed in the present study to avoid this problem. The process flow of the current PUREX in Japan and proposed process in the present study is shown in Fig. 5.102 [91]. If the fuel composed only of TRU like a Deep Burn fuel would be fabricated with the current PUREX without mixed denitration, the MOX product with high plutonium enrichment over 50 wt.% is generated and transported from reprocessing facilities to fuel fabrication facilities. The proliferation resistance becomes significantly depleted. Therefore the innovative LWR spent fuel reprocessing and fuel fabrication system was proposed as follows: The purified plutonium of nitrate solution is mixed with inert matrix of yttria-stabilized

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Figure 5.102 Pu flow in the reprocessing facilities [91].

zirconia (YSZ) instead of recovered uranium, and the route of the nitrate plutonium solution is connected to the fuel fabrication process of Sol-Gel method [92] directly. Due to the changes, the proliferation resistance is improved. This system can be established only by combining the current process of PUREX and the HTGR fuel fabrication without significant changes. The produced Clean Burn fuel, which is composed of the fuel compact matrix and TRi-ISOtropic (TRISO) fuel with the IMF kernel, has higher proliferation resistance than the LWR-MOX fuel. Here, the concept of conversion time defined by IAEA is referred. The definition of the conversion time [93] is the time required to convert different forms of nuclear material to the metallic components of a nuclear explosive device. It is one factor used to establish the timeliness component of the IAEA inspection goal via another concept of detection time. The material, which has a long conversion time, has high proliferation resistance. To reprocess the TRISO fuel, the complex head-end process, which surely increases the conversion time in comparison with the LWR reprocessing performing just chopping, is needed to connect the PUREX process. The head-end process developed in JAEA [94] is described as follows: First, the fuel compact matrix and outer PyC layer composed of graphite are removed by burning. Second, to remove SiC layer from the TRISO fuel particles, the layer is crashed mechanically by jet grind method or hard disk crusher. Third, to remove FP gas to reduce the radioactivity release for later processes, the fuel particles are burned again to adapt voloxidation method. This process is essential to remove inner PyC layer and buffer layer of fuel particles. There are great difficulties in the engineering of these processes especially for the crashing of SiC layer. SiC layer should be crashed effectively without crash of fuel kernel, which causes radioactive pollution. The exposure contributes the proliferation resistance, which is suggested

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by the difference of the conversion times between unirradiated fuel (order of weeks (13) and irradiated fuel (order of months (13)) [93]. Moreover, the IMF kernel of YSZ cannot be solved in ordinary nitrate solution owing to chemically inert characteristics [95]. To solve the YSZ kernel, it is inevitable to add the hydrogen fluoride [96]. The hydrogen fluoride is difficult to treat because that also corrodes the dissolver due to the very strong acids. As a result, it is found that the Clean Burn fuel has more “proliferation technical difficulty,” “proliferation cost,” and “proliferation time,” which are categorized as the essence of proliferation resistance in the Generation IV International Forum (GIF) than the LWRMOX fuel. This conclusion does not change even if innovative technology would be employed to treat Clean Burn spent fuel easily. This is because the innovative technology itself is regarded as “proliferation technical difficulty” and/or “proliferation cost,” which are the elements of proliferation resistance. As described above, with employing the proposed reprocessing and fuel fabrication system and IMF, the Clean Burn can incinerate plutonium without mixing with other actinoid nuclides with high proliferation resistance.

5.10.3 Core design Clean Burn has been developed based on the GTHTR300 to take advantage of the HTGR development experience in JAEA. However, the number of fuel columns is increased from 90 to 144, which is the same as Deep Burn as a result of reduction of the plutonium inventory per one fuel column to moderate neutron effectively. The plutonium nuclides have large neutron absorption cross-section compared with the uranium nuclides. If a large amount of plutonium is loaded in a few fuel columns, neutrons are absorbed before moderation to thermal energy region. Therefore the number of fuel columns should be increased. The core geometry is shown in Fig. 5.103 [97]. The fuel columns are composed of eight layer fuel blocks, which have a height of 1 m, for axial direction, and the core height is 8 m. The eight fuel layers are divided into four batches and shuffled when the fuels are reloaded. Two loading patterns are considered for fuel loading, that is, “sandwich loading” and “inout loading,” in the present study. The sandwich loading was invented for two-batch core in the previous GTHTR300 design study. The old fuels are sandwiched between new fuels. In the present study, the sandwich loading is extended for four-batch core. Inout loading is also employed considering neutronic economy. New fuels are placed at an inner region and old fuels are placed at an outer region. The loading patterns are shown in Fig. 5.104 [98]. The number of fuel columns is increased from 90 to 144 as same as Deep Burn. Moreover, plutonium inventory per fuel particle and packing fraction of CFP should be determined. Of course, the plutonium inventory should be preserved for the core considering neutron moderation. The plutonium inventory is determined to maintain the plutonium consumption ability with reasonable cycle length in the feasibility study on Deep Burn. But the IMF composition of PuO2-YSZ for the Clean Burn fuel should be determined. The mole fraction of PuO2 and YSZ can be determined freely according to design requests because both of them have the same crystal

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Figure 5.103 Core geometry of Clean Burn [97].

structure of fluorite. Thus in the this design, the mole fraction of PuO2 is varied from 20% to 50% as shown in Table 5.23 [99]. The fuel specification and calculation conditions, which refer to values for Deep Burn, are listed in Table 5.24 [76]. The discharged burnup of 500 GWd/t is set as target in this study by referring the Deep Burn design value of 546 GWd/t. The effect of the mole fraction of PuO2 to criticality is shown in Fig. 5.105. The achievable burnups of the one-batch core are almost the same value, which is 370 GWd/t. The burnups correspond to 590 GWd/t for the four-batch core. The target burnup of about 500 GWd/t is achievable. In addition, the Doppler coefficients are evaluated as shown in Fig. 5.106. The Doppler coefficients are defined as the reactivity gradient to fuel temperature from the operation temperature to that increasing with 100 C. Those coefficients are negative during the operation for all the cases. At the beginning of burnup, the coefficients show larger values for smaller mole fraction of PuO2. On the contrary, those show smaller values for larger mole fraction of PuO2 at the end of burnup. The mechanism is same as mentioned for the criticality. The increase of the resonance neutron capture due to the weakening of double heterogeneity enhances the Doppler reactivity effect, which is caused by weakening of a selfshielding effect due to the Doppler broadening of resonance capture peak. At the end of burnup, the trend of the Doppler coefficients is reversed by the change of the TRU vector due to the conversion. The low plutonium mole fraction can reduce the excess

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Figure 5.104 Diagram of loading pattern [98].

Table 5.23 Fuel compositions and packing fractions of Clean Burn [99]. PuO2 mole fraction (%)

Composition

PuO2 density (g/cm3)

Packing fraction (%)

20 30 40 50

64Zr-17Y-19Pu 56Zr-15Y-29Pu 48Zr-13Y-39Pu 40Zr-11Y-49Pu

2.46 3.63 4.77 5.87

46.5 31.6 24.0 19.5

Note: Pu includes MAs.

reactivity due to the weakening of double heterogeneity effect instead of BP without reduction of achievable burnup. However, it increases the packing fraction. Finally, the plutonium mole fraction is determined 30% considering the achievable packing fraction of 30% from the view point of fabrication. The neutronic feasibility was validated for the design with the plutonium mole fraction of 30%. The criticality during operation is shown in Fig. 5.107. The cycle lengths for the sandwich loading and inout loading are 250 and 263 days, respectively, and those correspond to the burnups of 500 and 526 GWd/t. The 239Pu consumption ratios are about 95%, fissile nuclide consumption ratios are about 80%, and neptunium and the precursor consumption ratios are about 40%, for both loading patterns. The 239Pu consumption ratios satisfy the “Cleanliness” of 95% defined in the present study. The neptunium and the precursor, that is, nuclide in 4N 1 1 series

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Table 5.24 Fuel specification and calculation condition of Clean Burn [76]. Items

Design values

Thermal power (MWt) Heavy metal inventory (t) Discharge burnup (GWd/t) Number of columns () Number of batches () Number of blocks per columns () Number of fuel pins per columns () Fuel column height (cm) Inner radius of fuel compact (cm) Outer radius of fuel compact (cm) Packing friction (%) Kernel diameter (μm) Buffer thickness (μm) IPC thickness (μm) SiC thickness (μm) OPC thickness (μm) Kernel density (g/cm3)

600 1.2 500 144 4 8 57 105 0.45 1.2

a

a

300 150 35 35 40 a

See Table 5.2.3.

Figure 5.105 Achievable burnup for one-batch core of Clean Burn [99].

decay chain, significantly contribute to the toxicity release from repository due to the high solubility of 237Np and 229Th, which is the daughter of 237Np, to the groundwater. The discharge burnup of clean burn is significantly high of about 500 GWd/t, and that of the plutonium consumption ratio is about 95% as same as Deep Burn.

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High Temperature Gas-cooled Reactors

Figure 5.106 Doppler coefficients for one-batch core of Clean Burn [76].

Figure 5.107 Criticality for four-batch core of Clean Burn [76].

5.10.4 Future study plan To establish the Clean Burn concept from the view point of engineering, we have the following subjects that need research and development: G

G

G

G

G

Detailed nuclear design considering the fuel temperature for safety analysis, Safety analysis for transient event, Fuel design and fabrication of Clean Burn fuel, Dose estimation for Clean Burn spent fuel in geological repository, and Introduction scenario for the Clean Burn reactor.

The detailed nuclear design is important to assess engineering feasibility. Installation of control rod and control rod program will be determined under the condition of a maximum fuel temperature limitation. The safety analysis for transient events should be performed. For the fuel design, there are some severe conditions for integrity of the Clean Burn fuel, that is, high fluence irradiation, much FP gas release, high yield

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of platinum group nuclides, and silver from plutonium. The coating technology of PuYSZ fuel kernel, including ZrC layer, has not been established yet. The engineering feasibility of this process will be validated by mockup examination. Dose estimation for the Clean Burn spent fuel in geological repository is important not only for safety but also for public acceptance. The spent fuel of Clean Burn is basically planned to be disposed of directly. The TRISO fuel shows integrity for a long time over one million years. Even after one million years, the failure fraction of fuel particles is predicted to be approximately 0.01%. Moreover, the chemical stability of YSZ is also expected to contribute to the safety in the geological repository. Introduction scenario for the Clean Burn reactor should be investigated. The benefit of the introduction will be found out by estimating mass balance under the realistic scenario.

5.11

Advanced fuel technology for reduction of highlevel radioactive waste

5.11.1 Introduction The GTHTR300 has particular features such as significantly high burnup of approximately 120 GWd/t, high thermal efficiency around 50%, and pin-in-block type fuel. The pin-in-block type fuel was employed to reduce processed graphite volume in reprocessing. By applying the feature, the effective waste loading method for direct disposal is proposed. By taking into account these feature, the number of HLW canister generations and its repository footprint are evaluated by burnup fuel composition, thermal calculation, and criticality calculation in repository. As a result, it is found that the number of canisters and its repository footprint per electricity generation can be reduced by 60% compared with the LWR representative case for direct disposal because of the higher burnup, higher thermal efficiency, less TRU generation, and effective waste loading proposed for HTGR [100]. But the reduced ratios change to 20% and 50% if the long-term durability of LWR canister is guaranteed. For disposal with reprocessing, the number of canisters and its repository footprint per electricity generation can be reduced by 30% compared with LWR because of the 30% higher thermal efficiency of HTGR.

5.11.2 Calculation for repository design The scenario, repository design, and specifications for disposal of a pressurized water reactor (PWR) are referred from the report of Japan Atomic Energy Commission (JAEC) [101]. According to this plan, the SFs are reprocessed after 4 years from discharge, and the vitrified wastes are disposed after 50 years from reprocessing (after 54 years from discharge). Also for direct disposal, the spent fuels are disposed after 54 years to match the plan with reprocessing. There are two parameters to determine the repository design: tunnel interval and waste package pitch. There are two limitations for those parameters from the safety requirement of structural integrity and maintenance of buffer function. For the first

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one, the limitations were evaluated by structural analysis. For the second one, the maximum temperature in bentonite buffer is problematic. The buffer function that delays nuclide migration will be lost when the temperature exceeds 100 C by thermal change of its property. The target value is set to be 90 C with considering uncertainties [101]. The maximum temperature of bentonite is evaluated by timedependent thermal transfer calculations for the HTGR cases using the ANSYS code, which solves the thermal equation by the finite element method with implicit time integral technique. The burnup composition and the decay heat are evaluated by using the ORIGEN code [102] for PWR and HTGR. The ORIGEN code is a burnup code for many purposes, such as evaluation of the source of radioactive waste for storage and disposal, and investigation on reactor strategy. The calculation conditions are listed in Table 5.25. The burnup of HTGR is approximately 120 GWd/t, and it is almost 3 times larger than PWR’s burnup of 45 GWd/t although the burnup days are almost same as that of PWR. In the same manner, specific power of HTGR is also larger than PWR’s. The fuel enrichment of 14 wt.% for HTGR is 3 times larger than PWR’s 4.5 wt.% because it should be proportional to the heat generation. The fuel compositions are listed in Tables 5.26 and 5.27. For the compositions at 54 year from discharge, those are evaluated for the spent fuels and reprocessed waste. The ratios where element is added into vitrified form are 0.442% and 0.548% for uranium and plutonium, respectively. For other actinoid element, the ratio of 100.0% is assumed [103]. The fuel compositions are normalized by initial heavy metal inventory and heat generation, respectively. The residual 235U inventory per initial heavy metal of 3.41 wt. % IHM is 3 times larger than PWRs of 1.12 wt.% IHM. (Here, IHM stands for initial heavy metal.) However, the value normalized heat generation of 0.285 kg/GWd is almost same as PWRs of 0.249 kg/GWd. The TRU nuclides inventory of HTGR of 0.151 kg/GWd is approximately half of PWRs of 0.273 kg/GWd in discharged fuel. For the reprocessing case, the amount of 0.022 kg/GWd is smaller than PWRs of 0.028 kg/GWd. TRU is composed of plutonium, minor actinoid (MA). MA includes neptunium, americium, and curium. Neptunium is generated from 235U by 2 times of neutron capture reactions and b-decay, and its generation is proportional to burnup because burnup is also reaction of 235U.

Table 5.25 Condition of burnup calculation [100].

Enrichment (wt.%) Specific power (MW/t) Burnup days (day) Burnup (GWd/t) Neutron flux at MOC (cm22/s) Capture cross-section of 238U at MOC (burn) Fission cross-section of 235U at MOC (burn) MOC, middle of cycle.

PWR

HTGR

4.5 38 1184.21 45 3.02 3 1014 0.88 33.68

14 84.63 1412.09 119.5 1.47 3 1014 3.28 59.30

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Table 5.26 Fuel compositions per initial heavy metal (wt.% IHM) [103]. PWR

234

U U 236 U 238 U 237 Np 238 Pu 239 Pu 240 Pu 241 Pu 242 Pu 241 Am 243 Am Total of U Total of TRU Total 235

Fresh

Discharged

At 54 years

At 54 years with reprocessing

0.00 4.50 0.00 95.50 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 100.00 0.00 100.0

0.00 1.12 0.57 92.43 0.06 0.03 0.63 0.25 0.18 0.25 0.01 0.01 94.12 1.23 95.37

0.01 1.12 0.57 92.43 0.07 0.02 0.64 0.25 0.01 0.25 0.16 0.01 94.13 1.23 95.37

0.00 0.00 0.00 0.41 0.07 0.00 0.00 0.01 0.00 0.01 0.03 0.01 0.42 0.12 0.54 HTGR

234

U U 236 U 238 U 237 Np 238 Pu 239 Pu 240 Pu 241 Pu 242 Pu 241 Am 243 Am Total of U Total of TRU Total 235

Fresh

Discharged

At 54 years

At 54 years with reprocessing

0.00 14.00 0.00 86.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 0.00 100.00 0.00

0.00 3.41 1.79 80.80 0.13 0.08 0.61 0.34 0.34 0.25 0.01 0.04 86.00 1.80

0.03 3.41 1.80 80.80 0.15 0.05 0.63 0.36 0.02 0.25 0.30 0.04 86.03 1.80

0.00 0.02 0.01 0.36 0.14 0.00 0.00 0.01 0.00 0.00 0.06 0.04 0.38 0.26

100.0

87.84

87.84

0.65

The decay heats per burnup of PWR and HTGR are shown in Fig. 5.108, and listed in Table 5.28 for actinoid nuclides showing major contribution for direct disposal and disposal with reprocessing. The decay heats of FPs completely coincide because the difference of fission yield does not give a significant difference [104]. For the case of direct disposal, the decay heat of actinoid nuclides of HTGR of 6.22 W/GWd is 20% lower

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High Temperature Gas-cooled Reactors

Table 5.27 Fuel compositions per burnup (kg/GWd) [100]. PWR

234

U U 236 U 238 U 237 Np 238 Pu 239 Pu 240 Pu 241 Pu 242 Pu 241 Am 243 Am Total of U Total of TRU Total 235

Fresh

Discharged

At 54 years

At 54 years with reprocessing

0.000 1.000 0.000 21.222 0.000 0.000 0.000 0.000 0.000 0.000 0.000 0.000 22.222 0.000 22.222

0.000 0.249 0.127 20.539 0.014 0.006 0.141 0.055 0.039 0.015 0.001 0.003 20.916 0.273 21.193

0.002 0.249 0.127 20.539 0.016 0.004 0.143 0.055 0.003 0.015 0.035 0.003 20.918 0.274 21.193

0.000 0.001 0.001 0.091 0.015 0.000 0.001 0.001 0.000 0.000 0.008 0.003 0.092 0.028 0.120

HTGR

234

U U 236 U 238 U 237 Np 238 Pu 239 Pu 240 Pu 241 Pu 242 Pu 241 Am 243 Am Total of U Total of TRU Total 235

Fresh

Discharged

At 54 years

At 54 years with reprocessing

0.000 1.172 0.000 7.197 0.000 0.000 0.000 0.000 0.000 0.000 0.000 0.000 8.368 0.000 8.368

0.000 0.285 0.150 6.761 0.011 0.006 0.051 0.029 0.028 0.021 0.001 0.004 7.197 0.151 7.351

0.002 0.285 0.150 6.761 0.013 0.004 0.052 0.030 0.002 0.021 0.025 0.004 7.200 0.151 7.351

0.000 0.001 0.001 0.030 0.012 0.000 0.000 0.001 0.000 0.000 0.005 0.004 0.032 0.022 0.054

than PWRs of 7.61 W/GWd at 54 years from discharged. Especially for 241Am, which shows the largest contribution, the decay heat of HTGR is 40% lower than PWRs. The major part of 241Am is generated by b-decay from 241Pu after discharge. For the case of reprocessing, the decay heat of plutonium, which is recovered in the reprocessing, is small, and the decay heat of 241Am, which is generated from 241Pu, is also small. Although, the decay heat of 241Am that is generated during operation and cooling time of

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Figure 5.108 Decay heat per burnup of PWR and HTGR. Rep stands for with reprocessing [100].

Table 5.28 Decay heat per burnup of major actinoid nuclides at 54 years from discharge (W/GWd) [100]. PWR

238

Pu Pu 240 Pu 241 Am 243 Am 244 Cm Total 239

HTGR

Direct disposal

Disposal with reprocessing

Direct disposal

Disposal with reprocessing

2.42 0.27 0.39 4.01 0.02 0.46 7.61

0.02 0.00 0.01 0.86 0.02 0.46 1.39

2.53 0.10 0.21 2.89 0.02 0.44 6.22

0.02 0.00 0.01 0.61 0.02 0.44 1.11

4 years until reprocessing is dominant, and the decay heat of HTGR of 0.61 W/GWd is 30% smaller than PWRs of 0.86 W/GWd. That of 244Cm for HTGR of 0.44 W/GWd is the almost same as PWRs of 0.46 W/GWd. The total value of HTGR of 1.11 W/GWd is 20% smaller than PWRs of 1.39 W/GWd.

5.11.3 Evaluation of waste package The feature of pin-in-block-type fuel of HTGR may be suitable to reduce the HLW volume, and the high burnup definitely contributes to reduce HLW volume. Then, the number of waste package generation is evaluated with considering these features in this study. The canister is assumed same as PWR based on KBS-3 concept by JAEC, but the material is carbon steel. The inner diameter is 864 mm, and the height is 4300 mm. Here, we plan to withdraw fuel rods from the fuel blocks like the case with reprocessing [93], and insert the fuel rods into the canister as shown in Fig. 5.109. The fuel rod

442

High Temperature Gas-cooled Reactors

Figure 5.109 Proposed waste loading method for HTGR and its criticality model in repository [100].

Figure 5.110 PWR canister and criticality model [100].

diameter is 26 mm, and the fuel rod length is 1050 mm. The fuel rods can be contained 3700 rods per canister, 925 rods per layer, and 4 layers per canister. The fuel rods correspond to 9% of the HTGR core, which has 41,040 fuel rods. Fig. 5.110 shows a canister of PWR and its criticality model, which assumed that the canister made of carbon steel and structural material of fuel assembly made

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of zircaloy such as cladding and grid spacer are completely corroded and flow out to outside of the buffer region. This assumption is realistic because the canister thickness is determined to endure the corrosion in the groundwater in 1000 years [103], and the structure material will be dissolved in 7600 years [105]. On the other hand, the fuel pellet spends a million years to be dissolved into the groundwater. After 10,000 years, the region inside the buffer region is filled with groundwater and fuel pellets [105]. The pellets distribute to realize the optimum moderation (moderator to fuel volume ratio is 3) from the viewpoint of criticality in the model. The footprint per canister was evaluated for HTGR and PWR. The tunnel interval is 32 m, and the waste package pitch is 10 m for PWR as same as reported by JAEC [105]. The tunnel interval is reduced to 24 m, and the waste package pitch is 8.5 m for HTGR with considering the heat generation. The interval and pitch are determined according to the structural limitations of 23.66 m and 8.232 m as reported by JAEC [105]. The results are shown in Fig. 5.111. The maximum bentonite temperature of 89.8 C for PWR shows a good agreement with that of 90 C reported by JAEC. The maximum bentonite temperature of HTGR is 85.6 C, and there is a margin of approximately 4 C even though the footprint is reduced. The footprint of PWR with the case of 4 assemblies per canister is 320 m2/canister. That of HTGR is 204 m2/canister. In addition, that of PWR with the case of 2 assemblies per canister of 192 m2/canister, which is almost the same as that of HTGR, but the number of canisters becomes twice. The number of canister generation for direct disposal and disposal with reprocessing was evaluated and compared with those of PWR. For direct disposal, the footprint in the repository is also evaluated with thermal calculations. The result is listed in Table 5.29 with the major specifications of reactor and fuel. The burnup of HTGR is almost 3 times higher than PWRs, and the thermal efficiency of HTGR is 30 higher than PWRs. Owing to the higher burnup, higher thermal efficiency, and the effective waste loading method, the number of canister generation of 1.20 canister/TWeh for HTGR is the lowest compared with the 2 assemblies and

Figure 5.111 Maximum bentonite temperature during disposal [100].

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High Temperature Gas-cooled Reactors

Table 5.29 Calculated specifications of waste package for direct disposal and disposal with reprocessing [100]. PWR

HTGR

45 34.5 2.684

119.5 45.6 0.765

Enrichment (wt.%) Burnup (GWd/t) Thermal efficiency (%) Heavy metal inventory per electricity generation (t/TWeh) Direct disposal Heavy metal inventory per canister (t/canister) Number of canister per electricity generation (canister/TWeh) Repository footprint per canister (m2/canister) Repository footprint per electricity generation (t/TWeh)

2 assemblies in a canister 0.920

4 assemblies in a canister 1.840

0.639

2.92

1.46

1.20

192

320

204

560.1

466.8

244.0

0.790

0.330

3.40

2.32

90

90

305.8

208.5

Disposal with reprocessing Heavy metal inventory per canister (t/canister) Number of canister per electricity generation (canister/TWeh) Repository footprint per canister (m2/canister) Repository footprint per electricity generation (t/TWeh)

4 assemblies per canister cases of PWR for direct disposal. It is reduced by 60% from 2 assemblies per canister case of PWR of 2.92 canister/TWeh, and reduced by 20% from 4 assemblies per canister case of PWR of 1.46 canister/TWeh. The footprint per electricity generation of 244.0 m2/TWeh for HTGR is the lowest compared with PWR’s cases. It is reduced by 60% from 2 assemblies per canister case of PWR of 560.1 m2/TWeh, and reduced by 50% from 4 assemblies per canister case of PWR of 466.8 m2/TWeh. The effective reduction of footprint per electricity generation is also owing to small footprint per canister of HTGR because of the low decay heat with the less TRU generation. That of 204 m2/canister for HTGR is almost same as 2 assemblies per canister case of PWR of 192 m2/canister. Comparing with the case of 4 assemblies per canister of 320 m2/canister, it is reduced by 40%. Subcriticality of repository is sufficiently remained for the direct disposal of HTGR. However, 4 assemblies per canister case of PWR achieve criticality without considering poison effect of FPs. In Japan, 2 assemblies per canister case should be the representative case of PWR with considering criticality safety in

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repository. On the other hands, 4 assemblies per canister case of PWR can be acceptable if the long-term durability of canister is guaranteed. For the disposal with reprocessing, the number of canister generation and repository footprint per electricity generation of HTGR are reduced by 30% because of the 30% higher thermal efficiency. The number of canister generation is determined by the heat generation limitation at vitrified form fabrication, and a major part of decay heat is generated from FPs, which is proportional to burnup. The number per electricity generation is reduced with the higher thermal efficiency. The repository footprint per canister is determined by the structural limitation, and does not depend on heat generation. If other disposal ways such as horizontal emplacement based on the KBS-3H concept are employed and/or scenario of reprocessing and disposal is changed, less TRU generation of HTGR contributes to reduce the footprint in repository.

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[94] J. Sumita, et al., Reprocessing technologies of the high temperature gas-cooled reactor (HTGR) fuel, Trans. AESJ 2 (2003) 546554. [95] H. Akie, et al., A new fuel material for once-through weapons plutonium burning, Nucl. Technol. 107 (1994) 182192. [96] N. Shirasu, et al., Burnup Measurement of Irradiated Rock-Like Fuels, JAERIResearch 2001-018, 2001. [97] M. Goto, et al., Development of security and safety fuel for Pu-burner HTGR (2) design study of fuel and reactor core, in: Proceedings of the 25th International Conference on Nuclear Engineering, Shanghai, China, ICONE25-67110, 2017. [98] Y. Fukaya, et al., Development on Nuclear Design Model for Detailed Design of Clean Burn HTGR, Japan Atomic Energy Agency, JAEA-Technology 2015-017, 2015. [99] Y. Fukaya, et al., Proceedings of the Series Presentations on Research and Development of Clean Burn HTGR—System Concept and Future Prospects, Japan Atomic Energy Agency, JAEA-Review 2014-010, 2014. [100] Y. Fukaya, et al., Reduction on high level radioactive waste volume and geological repository footprint with high burn-up and high thermal efficiency of HTGR, Nucl. Eng. Des. 307 (2016) 188196. [101] Japan Atomic Energy Commission, Report on Comparison of Nuclear Fuel Cycle Cost in the Basic Scenario, 2004. [102] G. Croff, ORIGEN2: a versatile computer code for calculating the nuclide compositions and characteristics of nuclear material, Nucl. Technol. 62 (1983) 335352. [103] Japan Nuclear Cycle Development Institute (JNC), Second Progress Report on Research and Development for the Geological Disposal of HLW in Japan; H12: Project to Establish the Scientific and Technological Basis for HLW Disposal in Japan, Project Overview Report, JNC TN1410 2000-001, 2000. [104] Y. Fukaya, et al., Investigation on spent fuel characteristics of Reduced-Moderation Water Reactor (RMWR)2008 Nucl. Eng. Des. 238 (2008) 16011611. [105] Japan Atomic Energy Commission (JAEC), Report on Comparison of Nuclear Fuel Cycle Cost in the Basic Scenario, 2004, URL: ,http://www.aec.go.jp/jicst/NC/iinkai/ teirei/siryo2004/kettei/sakutei041124.pdf..

Index

Note: Page numbers followed by “f” and “t” refer to figures and tables, respectively. A Abnormal events, 354 grouping of, 354 Acceptance criteria, 355 accidents, 355 anticipated operational occurrences, 355 HTTR, 165 167 Accident sequence grouping of, 354 identification of, 354 significant, 355 Accidents, 355 ACCORD code, 303 304 Advanced fuel technology for high burnup, 421 429 for plutonium burner, 429 437 for reduction of high-level radioactive waste, 437 445 AESJ. See Atomic Energy Society of Japan (AESJ) Air cooler, 112 Air ingress, 230 252 during primary pipe rupture accident, 399 421 phenomena during horizontal pipe break accident, 415 421 in reverse U-shaped channel, 399 415 process, 31 in reverse U-shaped channel, 233 243 in simulated reactor apparatus, 243 252 Alloy 800H, 119 120 Amoeba effect, 182 Analytical method, HTTR, 34 35 Anticipated occasional occurrences. See Anticipated operational occurrences (AOOs) Anticipated operational occurrences (AOOs), 21, 269, 352, 355, 424

Anticipated transients without scram (ATWS), 21 AOOs. See Anticipated operational occurrences (AOOs) Arbeitsgemeinschaft Versuchsreaktor (AVR), 10 As-fabricated fuel failure fraction, 181 182 Atomic Energy Society of Japan (AESJ), 349 350, 350f ATWS. See Anticipated transients without scram (ATWS) Auxiliary cooling system, 28, 113 Auxiliary helium systems, 151 155 AVR. See Arbeitsgemeinschaft Versuchsreaktor (AVR) Axial power distribution, 40 B Basic design earthquake ground motions, 94 Bayonet-type heat exchanger, 373 374 Beginning of life (BOL), 295 296 Bench-scale test, 371, 372f BFG. See Blast furnace gas (BFG) BFS system. See Blast furnace steelmaking system (BFS system) Blasius equation, 193 Blast furnace gas (BFG), 341 343 Blast furnace steelmaking system (BFS system), 341 Block-type HTGRs, 287 BLOOST-J2 code, 172 BOL. See Beginning of life (BOL) BPs. See Burnable poisons (BPs) Brayton cycle, 321 322 Brayton power conversion system, 337 Brine heater, 337 338 Buckling limit, 53 Bunsen reactor, 372 373

452

Burnable poison rods, 36 37 Burnable poisons (BPs), 32, 287, 315 317 Burnup, 34, 40 41 C Carbon capture and storage (CCS), 346 347 Carbon dioxide (CO2), 231 232 emissions, 341, 346, 346t Carbon monoxide (CO), 231 232, 422 Carbon structure in HTTR, 48 50 CBS. See Core bottom structure (CBS) CCS. See Carbon capture and storage (CCS) Ceramic heat exchanger, 372 Ceramic materials, 385 Ceramic-coated fuel particle, 314 315 CFD. See Computational fluid dynamics (CFD) CFPs. See Coated fuel particles (CFPs) Characteristic test of initial core, 269 278. See also Performance test; Safety demonstration test control rod characteristics, 274 275 critical approach, 271 272 excess reactivity, 272 273 neutron flux distribution, 276 278 nuclear calculations, 270 271 power distribution, 278 reactivity coefficient, 275 276 shutdown margin, 273 274 test condition, 270 CITATION-1000VP code, 34 Clean Burn, 314, 421, 429 430 criticality for four-batch core of, 436f fuel compositions and packing fractions of, 434t fuel fabrication process of, 430 432 fuel specification and calculation condition of, 435t CO. See Carbon monoxide (CO) Coal, 341 343 Coated fuel particles (CFPs), 2, 21, 180, 287, 422 Coaxial hot gas duct, 336 COG. See Coke oven gas (COG) Cogeneration, system design for, 329 340 HTGR renewable hybrid system, 338 340 hydrogen cogeneration, 330 336 seawater desalination, 336 338

Index

Coke oven (COv), 341 343 Coke oven gas (COG), 341 343 Combustible gas explosion, 396 397 coaxial pipe, 396f HTIVs, 397 399 Commercial HTGR, 313 314 Commissioning tests, 146 151 COMPARE-MOD1 code, 172 Compressor aerodynamic performance test, 359 Computational fluid dynamics (CFD), 399 Containment structure, 28, 138 151 emergency air purification system, 147 151 reactor containment vessel, 138 145 service area, 146 147 Control reactivity, 352 Control rods (CRs), 71, 114 126, 287 characteristics, 274 275 drive mechanism, 117 position instrumentation, 129 Control system, 29, 132 134 Control technology, 389 393 cooling system for helium with steam generator, 391f experimental and numerical results of system controllability test, 392f flow diagram of mock-up test facility for development of, 389f steam generator of mock-up test facility, 390f Controllable reactivity and shutdown margin, 37 38 Conventional modal decomposition method, 368 369 Coolant, 3 Coolant flow reduction test, 307 308 Cooling system, 102 113 Core bottom structure (CBS), 26, 101 102, 179 insulation performance of, 200 Core components, 25, 71 88, 191 203 vibration characteristics, 99 Core design, 432 435, 433f Core differential pressure instrumentation, 129 Core support graphite structures, 79 84 metallic structures, 85 87 plate, 264 268

Index

Core thermal-hydraulics, 41 48 design details, 42 44 design requirements, 41 42 evaluation results of design, 44 45 reevaluation of maximum fuel temperature, 45 48 Corrosion resistance, 384 385, 384f Coupling to heat application system, 352 353 COv. See Coke oven (COv) 1Cr 0.5Mo V steel, 70 21/4Cr 1Mo steel, 69 70 Creep analysis method, 68 buckling, 68 collapse, 218 220 fatigue, 220 223 damage, 68 interaction, 66 67 property, 65 rupture strength, 68 CRs. See Control rods (CRs) D DCHX. See Direct-contact heat exchanger (DCHX) Decarburization, 343 Deep Burn, 432 DELIGHT code, 34, 270 271 Depressurized flash concentration, 381 382 Desalination plant, 337 Deuterium, 394, 395f Diffusion coefficients, 406 Dioctyl phthalate (DOP), 150 151 Direct reduced iron (DRI), 343 Direct-contact heat exchanger (DCHX), 381 382 DOE. See U.S. Department of Energy (DOE) DOP. See Dioctyl phthalate (DOP) Doppler coefficient, 38 39, 433 434, 436f Doppler effect, 14 15 DRI. See Direct reduced iron (DRI) E EAF. See Electric arc furnace (EAF) EBU. See Extended burnup (EBU) ECT. See Eddy current testing (ECT) Eddy current testing (ECT), 228 229

453

Effective full power days (EFPD), 287, 426 427 Electric arc furnace (EAF), 341 Electricity consumption, 343 electricity hydrogen cogeneration, 343 generation thermal efficiency, 343 Electro-electrodialysis (EED) cell, 379 380 stack, 372 Electron-probe microanalyzer (EPMA), 182 183 Emergency air purification system, 147 151 End of life (EOL), 38 39 Engineered safety features actuating system, 135 systems, 28 Engineering Research Association of Nuclear Steelmaking (ERANS), 11 12, 172 EOL. See End of life (EOL) EPMA. See Electron-probe microanalyzer (EPMA) ERANS. See Engineering Research Association of Nuclear Steelmaking (ERANS) Excess reactivity, 32, 35 38, 272 273 Explosion of combustible gas, 396 397 Extended burnup (EBU), 422 F Fabrication technologies for HTTR fuel, 185 188 Failure mode effect analysis (FMEA), 354 Federal Republic of Germany (FRG), 10 Filter efficiency measurement, 150 151 First criticality, 257 258, 270 272 First phase test, 22 23 Fission products (FPs), 2, 21, 287, 422 424 FLOWGR code, 243 FLOWNET code, 42 43, 43f FLOWNET/TRUMP code, 173 Fluid dynamics, 31 FMEA. See Failure mode effect analysis (FMEA) Fort St. Vrain (FSV), 8 10 FPs. See Fission products (FPs) Fracture theory, 51

454

FRG. See Federal Republic of Germany (FRG) FSV. See Fort St. Vrain (FSV) Fuel, 1 3, 29 30, 71 73, 180 191 assembly of HTTR, 25, 26f blocks, 319 design, 424 427 fabrication process of clean burn, 430 432 technologies for HTTR fuel, 185 188 failure detection system, 129 131 performance of HTTR fuel, 189 191 research and development for design, 181 185 system, 155 157 handling system, 155 157 storage system, 157 Fukushima-Daiichi Nuclear Power Plant Accident, 421 Full power operation, prediction and test results at, 264 Full-scale mock-up model, 220 221 Full-scale outer tube model, 385 386 G GA. See General Atomics (GA) GAD. See Guidelines for Aseismic Design (GAD) Gamma-compensated ionization chambers (Gamma-CICs), 30 31 Gas circulators trip test, 307 308 Gas turbine, 313 314 power conversion system, 315 for power generation, 346, 359 370 Gas turbine generator (GTG), 319 Gas turbine modular helium reactor (GTMHR), 10 11 General Atomics (GA), 8, 429 430 Generation IV International Forum (GIF), 11, 432 Generator, 366 GIF. See Generation IV International Forum (GIF) Glass-lined components, 384 GRACE code, 172 Graphite, 1 3, 5 6, 30 components, 48 58 carbon structure in HTTR, 48 50 graphite design criteria, 50 57

Index

in-core graphite, 48 50 quality control, 58 structural integrity of, 100 102 oxidation, 31 GT-MHR. See Gas turbine modular helium reactor (GT-MHR) GTG. See Gas turbine generator (GTG) GTHTR300, 319 328 comparison of costs between GTHTR300 and PWR, 330f cross-sectional views of, 321f cycle diagram, 323f elements of high-performance helium gas compressor design approach, 327f gas turbine view, 323f heat exchanger module, 328f RPV cooling flow driven by intrinsic pressure gradients, 322f specifications, 320t of GTHTR300 compressor, 326t of GTHTR300 turbine, 325t stage-wise aerodynamic losses of helium gas turbines, 325f VCS, 329f GTHTR300C system, 314, 411, 415 416 apparatus simulated horizontal pipe rupture accident in, 415f Guidelines for Aseismic Design (GAD), 89 H Hardening ratio, 64 Hastelloy XR, 63 68, 334 development, 59 61 Heat application, 3 4 system, 387 399 coupling, 352 353 safety design for connection, 349 359 Heat balance, 343 of reactor cooling system, 280 281 Heat exchanger (HTX), 319 Heat removal from core, 352 Heat transfer, 31, 194 195 Heat transport piping design, 336 Helically coiled counter-type heat exchanger, 317 Helium (He), 232 coolant, 321 322 injection process, 399 purification system, 151 153

Index

sampling system, 154 storage and supply system, 154 155 Helium engineering demonstration loop (HENDEL), 14, 31, 121 122, 179, 191 Helium gas, 1, 3 4, 410 compressor, 313 314, 326, 359 360 design parameters for full and test compressors, 362t experimental and numerical results of axial velocity, 364f experimental and numerical results of performance, 364t helium compressor test rig, 363f helium test compressor in one-third dimensional scale, 361f helium gas-cooled synchronous type generator, 366 mixing performance of, 199 200 sealing performance of, 197 199 turbine, 357 359 HENDEL. See Helium engineering demonstration loop (HENDEL) Hexagonal graphite blocks, 74 79 HHT. See High temperature helium turbine (HHT) HHV. See Higher heating value (HHV) HI. See Hydrogen iodide (HI) High burnup, advanced fuel technology for, 421 429 design of high burnup fuel, 422 424 future study plan, 429 upgrade technologies for, 424 429 High quality steel (HQS), 341 343 High temperature engineering test reactor (HTTR), 1, 2f, 17 18, 257 258, 313 auxiliary helium systems, 151 155 chemistry control, 151 classification of HTTR facilities, 91t cogeneration demonstration, 355 359 containment structures, 138 151 cooling system, 25f, 29, 102 113, 132 134 core components, 71 88 core thermal-hydraulics, 41 48 design features, 18 31 development of process heat application system, 23 engineered safety systems, 28

455

evaluation of reactor performance, 21 22 fuel system, 155 157 graphite components, 48 58 high temperature operation validation using, 292 298 history and future plan of, 19 23 instrumentation system, 28 29, 128 131 metallic components, 58 70 nuclear design, 32 41 operation characteristic test of initial core, 269 278 construction and operation, 258 268 high temperature operation, 287 298 performance test, 278 287 safety demonstration test, 298 310 temperature rise of core support plate, 264 268 temperature rise of primary upper shielding, 258 264 performance test results, 135 138 process instrumentation, 131 R&D programs for, 29 31 reactivity control system, 113 127 reactor cooling system, 27 28 core, 25 27 internals, 71 88 safety demonstration test, 22 23 design, 158 173 protection system, 29, 134 135 seismic design, 89 102 specification of, 23t High temperature fission counter-chambers (HTFCs), 30 31 High temperature gas-cooled reactor (HTGR), 1, 17 18, 313 314, 349 359, 387 399 adaptability to environment, 6 closed-cycle power generation system, 339 340 construction in world, 9t features of, 1 6 flow diagram of HTGR hydrogen production system, 388f fuels development by JAEA, 423t gas turbine power plant, 319 328 hydrogen production system, 388f

456

High temperature gas-cooled reactor (HTGR) (Continued) renewable hybrid energy system, 339 340, 339f renewable hybrid system, 338 340 dynamic behavior of HTGR cogeneration plant, 340f load-following operational strategy for HTGR cogeneration plant, 339f research and development history in Japan, 11 15 in world, 7 11 roadmap for safety standard establishment, 349 350 safety, 4 6 safety guides acceptance criteria, 355 evaluation items, 353, 353t LBE selection, 354 355 safety requirements, 351 353, 351t confinement of radionuclides, 352 control reactivity, 352 coupling to heat application system, 352 353 heat removal from core, 352 loss-of-offsite power, 352 safety approach, 351 steam cycle power plant, 315 319 structure and materials, 1 3 High temperature helium gas, 335 336 High temperature helium turbine (HHT), 359 High temperature isolation valves (HTIVs), 314, 397 399, 397t, 398f High temperature operation, 287 298 test results of long-term, 288 292 validation using HTTR burnup data, 292 298 High temperature reactor, 341 Higher heating value (HHV), 343 High-level radioactive waste (HLW), 314 High-level radioactive waste, reduction of, 437 445 calculation for repository design, 437 441 evaluation of waste package, 441 445 High-sensitive gamma-uncompensated ionization chambers (HSUICs), 30 31

Index

HKG. See Hochtemperatur-Kernkraftwerk GmbH (HKG) HLW. See High-level radioactive waste (HLW) Hochtemperatur-Kernkraftwerk GmbH (HKG), 10 Horizontal pipe break accident, 415 421 experimental apparatus, 416 417, 416f experimental method, 417 experimental results, 417 421 air mole fraction changes, 419f cooling-water temperature, 419f flow velocity changes at outer pipe, 420f Hot spot evaluation of, 213 215 factors, 43 44, 46 47 HQS. See High quality steel (HQS) HSUICs High-sensitive gammauncompensated ionization chambers (HSUICs) HTFCs. See High temperature fission counter-chambers (HTFCs) HTGR. See High temperature gas-cooled reactor (HTGR) HTIVs. See High temperature isolation valves (HTIVs) HTR50S, 315 319 cross-sectional views of HTR50S reactor, 317f specifications, 316t of HTR50S steam generator, 318t system configuration of, 316f system layout of HTR50S vessel cooling system, 319f HTTR. See High temperature engineering test reactor (HTTR) HTTR hydrogen production system (HTTR-H2), 389 390 HTTR-GT/H2 plant, 356, 357f demonstration items, test items, and design target for, 358t plant bird’s eye view, 357f specifications of, 358t HTTR-H2. See HTTR hydrogen production system (HTTR-H2) HTX. See Heat exchanger (HTX) Hydrogen, 394 separation membrane, 380 381 steelmaking, 341

Index

Hydrogen (H2), 370 Hydrogen cogeneration, 330 336 cross-sectional view of GTHTR300C-I IHX, 333f of GTHTR300C-II IHX, 334f of heat transport piping, 335f GTHTR300C reactor system, 331f specifications of GTHTR300C, 331t of GTHTR300C IHX, 332t types of heat transport piping design, 335f Hydrogen iodide (HI), 336, 371 decomposer, 336 decomposer with hydrogen separation membrane, 380 381 conversion of HI decomposition and H2 production, 381f Hydrogen production, 349 cost, 348 349, 348f efficiency, 379 383 analytical estimation of hydrogen production thermal efficiency, 381 383 electro-electrodialysis cell, 379 380 hydrogen iodide decomposer with hydrogen separation membrane, 380 381 IS process technology for, 370 387 plant, 336, 356 system, 396 test, 377 379, 377f, 377t, 378f thermal efficiency, 381 383 heat and electricity input and H2 production thermal efficiency, 383t I ICRP. See International Commission on Radiological Protection (ICRP) IHX. See Intermediate heat exchanger (IHX) IHX primary coolant flow rate control system, 133 IMF. See Inert matrix fuel (IMF) In-core graphite, 48 50 In-core temperature monitoring system, 131 In-service inspection (ISI), 71, 82 84 In-service inspection technology (IST), 228 229 of tube, 228 229

457

Industrial material component test, 372 376 Bayonet-type H2SO4 decomposer, 374f concentration change in catholyte and anolyte, 375f external-circulation gas liquid cocurrent-type Bunsen reactor, 373f O2 production rate against H2SO4 feed rate to decomposer, 374f radial-flow-type adiabatic fix-bed HI decomposer, 376f Inert matrix fuel (IMF), 314 Inherent safety features of HTGR, 4, 14 15 Instrumentation system, 28 29, 128 131 Insulation performance of CBS, 200 Integration technology, 13 14 Intermediate heat exchanger (IHX), 10, 20, 106 108, 106f, 179, 217 229, 330 331, 387 388 creep collapse, 218 220 creep fatigue, 220 223 in-service inspection technology of tube, 228 229 seismic behavior of tube bundle, 223 226 thermal hydraulic behavior of tube bundle, 226 228 International Commission on Radiological Protection (ICRP), 158 160 Iodine sulfur (IS) process technology for hydrogen production, 370 387 bench-scale test, 371 component materials, 383 387 corrosion resistance, 384 385 strength, 385 387, 386f elemental technologies, 371 372 hydrogen production test, 377 379 improvement of hydrogen production efficiency, 379 383 industrial material component test, 372 376 Iodine sulfur method (IS method), 19, 314, 370 371 IPyC. See Isotropic high-density carbon (IPyC) Iron ore, 341 343 Irradiation test, 427 429, 428f IS method. See Iodine sulfur method (IS method) ISI. See In-service inspection (ISI) Isolation valves, 356, 397

458

Isothermal condition, 237 239 Isotropic high-density carbon (IPyC), 180 IST. See In-service inspection technology (IST) J JAEC. See Japan Atomic Energy Commission (JAEC) Japan Atomic Energy Agency (JAEA), 1, 10 11, 371 300 MWe class nuclear helium gas turbine design, 360 361, 360f Japan Atomic Energy Commission (JAEC), 437 Japan Material Test Reactor (JMTR), 29 30, 131, 179, 428 429 Japan research reactor-2 (JRR-2), 29 30 Japanese Industrial Standards (JIS), 64 JMTR. See Japan Material Test Reactor (JMTR) JRR-2. See Japan research reactor-2 (JRR-2) K K-type thermocouples, 417 Kernel migration, 182 L LBEs. See Licensing basis events (LBEs) LDG. See Linz Donawitz converter gas (LDG) Leakage-rate test, 140 145 Licensing basis events (LBEs), 354 selection, 354 355 abnormal events and PIEs, 354 accident sequence, 354 grouping of abnormal events, 354 grouping of accident sequence, 354 safety function and mitigation system, 354 selection, 355 significant accident sequence, 355 for single failure events, 354 Light hydrocarbon, 12 13 Light water reactor (LWR), 1, 314, 351t, 352 Linz Donawitz converter gas (LDG), 341 343 Liquid metal fast breeder nuclear reactors (LMFBRs), 69

Index

Liquid liquid separation, 371 LMFBRs. See Liquid metal fast breeder nuclear reactors (LMFBRs) Loading patterns, 432, 434f LOCA. See Loss of coolant accident (LOCA) LOFC test. See Loss of forced cooling test (LOFC test) Loss of coolant accident (LOCA), 271 272 Loss of forced cooling test (LOFC test), 14 15, 21, 309 310 “Loss of off-site electric power”, 124 Loss-of-offsite power, 352 LWR. See Light water reactor (LWR) M MA. See Minor actinoid (MA) Magnetic bearings, 322, 366 370 actual magnetic bearing suspended rotor system, 367t development test, 359 360 equivalent mass distribution of numerical model, 368f multiinput multioutput control system, 367f numerical result of vibration mode of rotor, 368f effect of reduction method on unbalance response, 369f Mass balance, 343 of steelmaking systems, 344t Master logic diagram (MLD), 354 Membrane distillation process, 337 338 Metallic components, 58 70 design limits and rules, 61 70 Hastelloy XR development, 59 61 identification of failure modes, 61 Metallic materials, 30 MEXT. See Ministry of Education, Culture, Sports, Science and Technology (MEXT) MHTGRs. See Modular high temperature gas-cooled reactors (MHTGRs) Midrex steelmaking (MS), 341 MIMO controller. See Multiinput multioutput controller (MIMO controller) Ministry of Education, Culture, Sports, Science and Technology (MEXT), 15, 20 21

Index

Minor actinoid (MA), 438 Mitigation System (MS), 160 MLD. See Master logic diagram (MLD) Moderator, 3 Modified Coulomb Mohr theory, 54 Modular high temperature gas-cooled reactors (MHTGRs), 203 Molecular diffusion, 408 Monte Carlo simulation, 385 MS. See Midrex steelmaking (MS); Mitigation System (MS) MSF. See Multistage flash (MSF) Multicomponent gas, 406 Multiinput multioutput controller (MIMO controller), 313 314 Multistage flash (MSF), 313 N N-type TCs. See Nicrosil Nisil thermocouples (N-type TCs) N-type thermocouples, 28 29 Natural convective heat transfer, 211 213 Natural gas, 343 NDCS. See Nuclear design code system (NDCS) Neutron detector, 128 Neutron flux distribution, 276 278 monitoring system, 136 Next-generation nuclear plant (NGNP), 11, 399 NFI. See Nuclear Fuel Industries, Ltd. (NFI) NHS process. See Nuclear hydrogen steelmaking process (NHS process) Nicrosil Nisil thermocouples (N-type TCs), 30 31 Nitrogen (N2), 232 Nonflammable gas, 396 397 NRA. See Nuclear Regulation Authority (NRA) NS. See Nuclear steelmaking (NS) Nuclear cogeneration demonstration, 356 Nuclear design, HTTR, 32 41 analytical method, 34 35 design requirement, 32 34 evaluation of nuclear characteristics, 35 41 Nuclear design code system (NDCS), 34 Nuclear energy, 11, 387

459

Nuclear Fuel Industries, Ltd. (NFI), 180 Nuclear grade alloy Hastelloy XR, 61 Nuclear heat, 1, 13 14, 343 Nuclear hydrogen production cost, 341 Nuclear hydrogen steelmaking process (NHS process), 313, 341 Nuclear instrumentation, 128 129 Nuclear Regulation Authority (NRA), 14 15, 21, 356 Nuclear shutdown margin, 35 38 Nuclear steelmaking (NS), 3 Numerical analysis, 243 246 Numerical method, 208 213 O Oarai Gas Loop-1 (OGL-1), 14, 29 30 OECD. See Organization for Economic Cooperation and Development (OECD) 1/3-scale nominal design blade, 361 362 turbo-compressor, 360 Onset time of natural circulation, 413 415 flow through apparatus, 413 Operational mode selector, 132 OPyC. See Outer isotropic high-density carbon (OPyC) Organization for Economic Cooperation and Development (OECD), 8 ORIGEN code, 438 Outer isotropic high-density carbon (OPyC), 180 Oxidation effect, 56 57 OXIDE-3F code, 172 P Palladium SiC interaction, 182 183 Parallel-loaded operation, 132 Partial leakage-rate test, 143 Passive cooling system, 203 217 evaluation of hot spot, 213 215 evaluation of local hot spot, 215 217 experiment, 205 208 numerical method, 208 213 PBMR. See Pebble-bed modular reactor (PBMR) PCRV. See Prestressed concrete reactor vessel (PCRV) Pebble-bed modular reactor (PBMR), 10 11

460

Performance test, 278 287. See also Characteristic test of initial core; Safety demonstration test fuel and fission product behavior, 286 287 heat balance of reactor cooling system, 280 281 heat exchanger performance, 281 282 major test items, 279 280 reactor control system performance, 282 284 residual heat removal performance, 284 285 thermal expansion performance, 285 286 Permanent reflector block, 26 PIEs. See Postulated initiating events (PIEs) Pin-in-block-type fuel, 441 442 element, 2 3, 71, 429 Plant control device, 133 134 Plutonium (Pu), 421 advanced fuel for plutonium burner, 429 437 core design, 432 435 fuel fabrication process of Clean Burn, 430 432 future study plan, 436 437 inventory, 432 433 Pu-burner HTGR, 422 Plutonium dioxide (PuO2), 422 Postulated initiating events (PIEs), 354 Power coefficient, 38 39, 276 Power distribution, 34, 39 40, 278 Power generation, 1, 4, 6, 8 system design for, 314 328 HTR50S, 315 319 Power unbalance, 195 196 PPWC. See Primary-pressurized water cooler (PPWC) Preservice inspection (PSI), 83 Pressure drop, 193 194 Pressurized helium gas, 332 Pressurized water pump, 112 Pressurized water reactor (PWR), 328, 437 Pressurized water temperature control system, 134 Pressurized water-cooling system, 112 Prestressed concrete reactor vessel (PCRV), 8 10 Prevention System (PS), 160

Index

Primary concentric hot gas duct, 109 110 Primary cooling system, 103 110 Primary gas circulator, 108 109 Primary helium pressure control system, 134 Primary pipe rupture accident, 230 252 Primary upper shielding, 258 264 Primary water cooler (PWC), 199 Primary-pipe rupture, 31 Primary-pressurized water cooler (PPWC), 20, 103 105, 147 148, 280 281 primary coolant flow rate control system, 134, 136 137 Primary-pressurized water differential pressure control system, 134 Primary secondary helium differential pressure control system, 134 Probabilistic safety assessment (PSA), 396 397 Process heat application system, 23 Process heat supply, 349 Process instrumentation, 131 PS. See Prevention System (PS) PSA. See Probabilistic safety assessment (PSA) PSI. See Preservice inspection (PSI) PWC. See Primary water cooler (PWC) PWR. See Pressurized water reactor (PWR) Pyrolytic carbon (PyC), 8 R R&D. See Research and development (R&D) R/B. See Releases to birth ratio (R/B) Radial power distribution, 39 40 Radionuclides, 329 330 confinement of, 352 RAHP. See Research Association of High Temperature Gas Cooled Reactor Plant (RAHP) Rankine power conversion cycle, 336 337 RATSAM6 code, 172 Reactivity addition rate, 34, 38 coefficient, 34, 38 39, 275 276 control system, 26 27, 113 127 control rod system, 114 126 reserve shutdown system, 126 127 insertion test, 304 307 losses, 36

Index

Reactor containment vessel, 138 145 control system performance, 282 284 cooling system, 27 28 heat balance of, 280 281 core, 25 27 core components, 25 reactivity control system, 26 27 reactor internals, 26 RPV, 27 inlet coolant temperature control system, 133 instrumentation, 30 31, 128 131 internals, 26, 71 88, 191 203 outlet coolant temperature control system, 132 133, 137 138 physics, 30 power control device, 132 system, 132 protection system, 134 135 shutdown margin, 32 33 Reactor pressure vessel (RPV), 20, 27, 258, 315 Recirculation heat recovery section, 337 338 Recoverable waste heat (RWH), 337 Reducing gas heater (RGH), 12 13 Reheating cycle, 315 Releases to birth ratio (R/B), 183 184, 189 Reliability test of control rods, 121 122 Renewable energy power generation, 338 Repository design, calculation for, 437 441 decay heat per burnup of major actinoid nuclides, 441t Research and development (R&D), 1 on commercial HTGR advanced fuel for plutonium burner, 429 437 advanced fuel technology for high burnup, 421 429 advanced fuel technology for reduction of high-level radioactive waste, 437 445 gas turbine technology for power generation, 359 370 IS process technology for hydrogen production, 370 387

461

prevention technology for air ingress, 399 421 safety design for connection of heat application system and HTGR, 349 359 system design for cogeneration, 329 340 system design for power generation, 314 328 system design for steelmaking, 341 349 system integration technology for connection of heat application system, 387 399 on components, 179 180 air ingress, 230 252 core components, 191 203 fuel, 180 191 IHX, 217 229 passive cooling system, 203 217 reactor internals, 191 203 history in Japan, 11 15 in world, 7 11 of HTGR in Japan, 3 JAEA, 180 programs for HTTR, 29 31 related R&D for fuel design, 181 185 Research Association of High Temperature Gas Cooled Reactor Plant (RAHP), 349 350 Reserved shutdown system, 37 38, 126 127 Residual heat removal performance, 284 285 system, 113 Reverse osmosis (RO), 337 338, 381 382 Reverse U-shaped channel change in air concentration, 403f change in velocity of natural air circulation, 401f, 402f experimental apparatus, method, and results, 399 406, 400f numerical analysis, 406 412 analytical model for 1D analysis, 407f distribution of molar fraction of helium gas, 409f velocity change with time elapsing, 408f

462

Reverse U-shaped channel (Continued) onset time of natural circulation, 413 415 flow, 404t flow through apparatus, 413 wall and gas temperature distribution for, 400f Reverse U-shaped channel, 233 243 Reynolds number, 359, 365, 365f RGH. See Reducing gas heater (RGH) Rise-to-power test, 20 21, 135 RO. See Reverse osmosis (RO) Roadmap for safety standard establishment, 349 350, 350f Rod-type burnable poisons, 295 297 RPV. See Reactor pressure vessel (RPV) Ruhrstahl-Heraeus (RH) procedure, 341 343 RWH. See Recoverable waste heat (RWH) S Safety demonstration test, 22 23, 298 310. See also Characteristic test of initial core; Performance test analysis code and model, 303 304 coolant flow reduction test, 307 308 gas circulators trip test, 307 308 HTTR control system, 300 302 loss of forced cooling test, 309 310 reactivity insertion test, 304 307 safety demonstration test plan, 302 303 Safety design, 158 173 acceptance criteria, 165 167 philosophy, 158 160 safety classification, 160 163 safety criteria, 173 safety evaluation technologies, 169 173 safety functions, 164 165 selection of events, 167 169 Safety protection system, 29, 134 135 Schur decomposition, 368 369 Science and Technology Agency (STA), 19 Scram tests, 120 121 Sealing performance of helium gas, 197 199 Seawater desalination, 313, 336 338 intermediate water loop coupling of MSF to nuclear plant, 337f optimized incrementally loaded MSF process, 338f

Index

Second intermediate heat exchanger (Second IHX), 356 Secondary gas circulator, 112 Secondary helium cooling system, 110 112 Secondary helium piping, 112 Secondary pressurized water cooler (SPWC), 27 28, 112 Seismic behavior of tube bundle, 223 226 Seismic design, 89 102 basic design earthquake ground motions, 94 development of evaluation method, 99 geological composition, 94 guideline of, 90 seismic classification, 90 94 seismometry, 94 structural integrity of graphite components, 100 102 structure of core components, 94 98 Seismometry, 94 Service area, 146 147 SF. See Shaft furnace (SF) SG. See Steam generators (SG) SH. See Steam heater (SH) Shaft furnace (SF), 343 Shielding blocks, 87 88 Shock absorber, 116 Shutdown margin, 273 274 SiC. See Silicon carbide (SiC) Signal isolating auxiliary cooling water line, 135 containment vessel, 135 Signal starting up auxiliary cooling system, 135 Silicon carbide (SiC), 180, 372, 422 Single failure events, LBE selection for, 354 Single-crystal Ni-based alloy, 321 322 Single-input single-output controller (SISO controller), 369 370 Single-loaded operation, 132 SISO controller. See Single-input singleoutput controller (SISO controller) Small-scale outer tube model, 385 386 Solar photovoltaic, 339 SONATINA-2V code, 99 Spherical fuel, 1 SPWC. See Secondary pressurized water cooler (SPWC) SR. See Steam reformer (SR)

Index

STA. See Science and Technology Agency (STA) Start-up test, 147 150 Steam generators (SG), 12 13, 315, 317 Steam heater (SH), 12 13 Steam reformer (SR), 12 13, 390 Steelmaking, system design for, 341 349 CO2 emission, 346 cost, 346 349, 347f flow diagram of steelmaking systems, 341 345, 342f heat input to steelmaking systems, 345t mass balance of steelmaking systems, 344t Straight tube model, 385 386 Stress analysis, 54 55, 125 126 Stress classification, 51 52 Stress limit, 52 53 Structural integrity of graphite components, 100 102 Sulfur trioxide (SO3), 373 374 Sulfuric acid (H2SO4), 336, 371 decomposer, 314 Surface roughness, 365 366 Surveillance test, 82 84 SUS316 austenitic stainless steel, 70 SUS321TB austenitic stainless steel, 70 System design for cogeneration, 329 340 for power generation, 314 328 for steelmaking, 341 349 System integration technology, 387 399 control technology, 389 393 explosion of combustible gas, 396 397 tritium permeation, 393 396 T TAC-NC code, 172 Temperature analysis, 122 124 coefficient, 275 276 THANPACST2 code, 204, 208 Thermal absorber, 389 Thermal expansion performance, 285 286 Thermal hydraulic behavior of tube bundle, 226 228 characteristics, 192 195 Thermal performance of coaxial hot gas duct, 201 203

463

Thermal power plants, 336 337 Thermochemical H2 production, 370 371 Thermochemical water splitting process, 313, 329 330 IS process, 341 Thorium Hochtemperatur Reaktor (THTR300), 10 Three Mile Island Unit-2 (TMI-2), 1 THTR-300. See Thorium Hochtemperatur Reaktor (THTR-300) THYDE-HTGR code, 124, 172 TMI-2. See Three Mile Island Unit-2 (TMI-2) Trans uranium (TRU), 421 TRi-ISOtropic. See Tristructural isotropic (TRISO) Tristructural isotropic (TRISO), 180, 422 coating, 131 fuel, 431 432 Tritium permeation, 393 396, 394f deuterium, 394 hydrogen, 394 TRU. See Trans uranium (TRU) Tube bundle seismic behavior, 223 226 thermal hydraulic behavior, 226 228 Turbine bypass control, 339 340 Turbine cooling scheme, 324, 324f TWOTRAN code, 270 271 TWOTRAN-2 code, 34 U U.S. Department of Energy (DOE), 10 11 Upgrade technologies for high burnup fuel design, 424 427 irradiation test, 427 429 Uranium dioxide (UO2), 180 Uranium fuel, 343 V Very high temperature reactor (VHTR), 8, 11 14 Very high temperature reactor critical assembly (VHTRC), 14, 30, 35, 185 Vessel cooling system (VCS), 21, 28, 113, 315 VHTR. See Very high temperature reactor (VHTR)

464

Index

VHTRC. See Very high temperature reactor critical assembly (VHTRC) VIENUS code, 78 79

Weibull distribution, 425 Whole core burnup calculations, 297 298 Whole leakage-rate test, 143 145

W Waste package, 441 445, 444t maximum bentonite temperature during disposal, 443f PWR canister and criticality model, 442f waste loading method for HTGR, 442f

Y Yttria-stabilized zirconia (YSZ), 422, 430 431 Z Zirconium carbide (ZrC), 422