PROCEEDINGS OF THE
Sixteenth International Cryogenic Engineering Conference] International Cryogenic Materials Conference
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PROCEEDINGS OF THE
Sixteenth International Cryogenic Engineering Conference] International Cryogenic Materials Conference PART 1
Kitakyushu, Japan 20-24 May 1996
Editors: T. Haruyama, T. Mitsui and K. Yamafuji
ELSEVIER
U.K.
Elsevier Science Ltd, The Boulevard, Langford Lane, Kidlington, Oxford, OX5 1GB, England
U.S.A.
Elsevier Science Inc., 660 White Plains Road, Tarrytown, New York, 10591-5153, U.S.A.
JAPAN Elsevier Science Japan, Tsunashima Building Annex, 3-20-12 Yushima, Bunkyo-ku, Tokyo 113, Japan
Copyright © 1997 Elsevier Science All Rights Reserved. No part of this publication may be reproduced, stored in a retrieval system or transmitted in any form or by any means; electronic, electrostatic, magnetic tape, mechanical photocopying, recording or otherwise, without prior permission in writingfrom the publisher. First edition 1997 Library of Congress Cataloging in Publication Data A catalogue record for this tit!e is available from the Library of Congress. British Library Cataloguing in Publication Data A catalogue record for this title is available from the British Library.
ISBN 0-08-042688-3
Printed and bound in Great Britain by BPC Wheatons Ltd, Exeter
CONTENTS LIST
PART ONE Committee members Exhibitors Foreword Greetings from the ICE Committee Greetings from the ICMC Board
oo
XXXVll
xxxix xli xliii xlv
Mendelssohn Award
xlvii
Mendelssohn Award Winner
xlix
Plenary Lectures A way to commercialization of high-Tc superconductors S. Tanaka Development of the superconducting Maglev system in Japan M. Ozeki
13
Cryogenic properties of advanced composites and their applications S. Nishijima
19
122-Channel SQUID neuromagnetometer for studies of information processing in the human brain O. V. Lounasmaa
27
Advances in cryocoolers R. Radebaugh
33
The Large Hadron Collider project L. R. Evans
45
Overview of the ITER project R. A ymar
53
Large Scale Refrigeration Refrigeration Plants
Design and construction of cryogenic components for LHD S. Satoh, T. Mito, S. Yamada, J. Yamamoto, O. Motojima and LHD Group
63
vi
Contents
The constituent hardware of the large helium refrigeration plant for LHD H. Matsuda, I. Ushijima, M. Katada, S. Satoh, J. Yamamoto and O. Motojima
67
Simulation of the large helium refrigeration plant for LHD M. Nobutoki, K. Iwamoto and H. Matsuda
71
Construction report of 10 kW class helium refrigerator for LHD S. Satoh, T. Mito, S. Yamada, A. Iwamoto, R. Maekawa, S. Moriuchi, T. Baba, K. Ooba, H. Sekiguchi, J. Yamamoto, O. Motojima and LHD Group
75
Cryogenic control system for the large helical device T. Mito, S. Satoh, R. Maekawa, S. Yamada, K. Takahata, A. Iwamoto, H. Yamada, K. Watanabe, T. Baba, S. Moriuchi, K. Oba, H. Sekiguchi, K. Murai, K. Iimura, K. Nakamura, J. Yamamoto, O. Motojima and LHD Group
79
Liquefaction control of 10 kW class cryogenic system for the LHD S. Yamada, S. Satoh, T. Mito, R. Maekawa, A. Iwamoto, S. Moriuchi, T. Baba, J. Yamamoto, O. Motojima and LHD Group, H. Matsuda, I. Ushijima, K. Nakamura, T. Kukano and M. Katada
83
Study of optimum operating condition for the helium refrigeration system of the LHD with a dummy load apparatus R. Maekawa, T. Mito, S. Yamada, S. Satoh, A. Iwamoto, T. Baba, S. Moriuchi, K. Ohba, H. Sekiguchi, I. Ohtake, H. Yamada, J. Yamamoto, O. Motojima and LHD Group, K. Chida and T. Fukano
87
Cryogenic operation and testing of the extended LHC prototype magnet string A. BOzaguet, J. Casas-Cubillos, H. Guinaudeau, B. Hilbert, Ph. Lebrun, L. Serio, A. Suraci and R. van Weelderen
91
Demands in refrigeration capacity for the Large Hadron Collider Ph. Lebrun, G. Riddone, L. Tavian and U. Wagner
95
Simulation program for cryogenic plants at CERN E. Melaaen, G. Owren, A. Wadahl and U. Wagner
99
Operation of the cryogenic system for superconducting cavities in LEP M. Barranco-Luque, S. Claudet, Ph. Gayet, N. Solheim and G. Winkler
103
Thermodynamic booster for the CERN Omega cryoplant F. Haug, J.-P. Dauvergne, H. Rieder and P. Chaffard
107
Refrigeration system for the ATLAS experiment F. Haug, J. I.-P. Dauvergne, G. Passardi, D. Cragg, C. Cure, P. Pailler, C. Mavri and A. Yamamoto
111
The cryogenic system of the ATLAS experiment end cap toroids D. Cragg
115
Cryogenic design of the ATLAS thin superconducting solenoid magnet K. Tanaka, A. Yamamoto, Y. Doi, Y. Makida, T. Haruyama and T. Kondo
119
Barrel toroid cryogenic system for the ATLAS detector C. Mayri, C. Cur~, R. Duthil, D. Cragg, F. Haug and G. Passardi
123
Performance study of the cryogenic system for ITER CS model coil T. Kato, K. Hamada, K. Kawano, K. Matsui, T. Hiyama, K. Nishida, T. Honda, M. Taneda, S. Sekiguchi, K. Ootsu, H. Tsuji, S. Shimamoto, H. Yamamura, I. Kawashima, Y. Nakayama and Y. Watanabe
127
Contents vii Engineering design of cryoplant for ITER K. Hamada, T. Kato, T. Honda, K. Matsui, K. Nishida, H. Tsuji, S. Shimamoto, K. Yoshida, M. E. P. Wykes, V. Kalinin, M. Mori and A. Miyake
131
Interface detail design for ITER coil system T. Honda, F. Iida, K. Matsui, Y. Yasukawa, K. Nishida, K. Hamada, T. Kato, H. Tsuji, K. Yoshida, K. Sakaki, H. Hiue and S. Shimamoto
135
Development and operating experience of the nuclotron cryogenic system N. Agapov, H. Khodzhibagiyan, A. Kovalenko, A. Smirnov and A. Sukhanova
139
Mathematical modeling of a Fermilab helium liquefier coldbox M. G. Geynisman and R. J. Walker
143
HERA at lower temperatures?- Operational test of the HERA cryogenic system at subatmospheric pressure H. Lierl and H. Herzog
147
Cryogenics for tokamak HT-7 SC toroidal magnet Y.-F. Bi, N. Qiu and J.-R. Wang
151
Theoretical calculation of the large hydrogen liquefaction process K. Iwamoto
155
Development of helium refrigeration systems for 70 MW class superconducting generators H. Yanagi, M. Ikeuchi, A. Machida and Y. Ikeda
159
Long Term Operation Reliability of helium refrigeration systems for the TRISTAN detector magnets Y. Doi Technical analysis and statistics from long term helium cryoplant operation with experimental superconducting magnets at CERN D. Delikaris, J. L-P. Dauvergne and F. Haug
165
169
Cryogenics for CERN experiments. Past, present and future J. Bremer, J. I.-P. Dauvergne, D. Delikaris, N. Delruelle, F. Haug, G. Kesseler, G. Passardi, J. M. Rieubland and J. Tischhauser
173
Present state of Tevatron lower temperature operation B. L. Norris
179
Cryogenic system for TRISTAN superconducting RF cavities: Description and operating experience K. Hosoyama, K. Hara, A. Kabe, Y. Kojima, T. Ogitsu, Y. Morita, Y. Sakamoto, H. Nakai, T. Fujita and T. Kanekiyo
183
Compressors High power refrigeration at temperatures around 2.0 K G. Gistau-Baguer
189
A cryogenic axial-centrifugal compressor for superfluid helium refrigeration L. Decker, K. LO'hlein, P. Schustr, M. Vins, I. Brunovsk, L. Tt~cek, Ph. Lebrun and L. Tavian
195
Upgrade of the CERN cryogenic station for superfluid helium testing of prototype LHC superconducting magnets V. Benda, J. I.-P. Dauvergne, F. Haug, S. Knoops, Ph. Lebrun, F. Momal, V. Sergo, L. Tavian and B. Vullierme
199
viii
Contents
Performance analysis of multistage 80 K centrifugal compressors for helium refrigerator
203
H. Asakura, N. Saji, Y. Kaneko, S. Yoshinaga, M. Mori, J. Sato, A. Miyake, T. Iwasaki, I. Nishimura, T. Hosoya and T. Umeda
Design manufacture and consideration for test result of centrifugal cold compressor for TEVATRON lower temperature upgrade
207
N. Saji, Y. Kaneko and H. Asakura
Development of helium oil free screw compressor
211
K. Kitagawa, Y. Hirao, Y. Yanagi and Y. Ikeda
Turboexpanders The experimental study of self-acting gas-lubricated tilting-pad thrust bearings for cryogenic turboexpander
217
C.-Z. Chen, H. Yao and Y. Wu
A genetic algorithm based optimization design method for cryogenic turboexpander
221
C.-H. Gao, H. Yao, and C.-Z. Chen
Predicting performance of helium expansion turbines using similarity principles
225
L.-Q. Liu and C.-Z. Chen
Stability study of herringbone-grooved, journal gas bearing for small cryogenic expansion turbine
229
L.-Q. Liu, G.-L. Zhou and C.-Z. Chen
Development of foil journal bearing for small high speed cryogenic turboexpander
233
H. Yao, H.-Y. Quan and C.-Z. Chen
Cryocoolers Pulse Tube Coolers Evaluation of experimental pulse tube refrigerator data with predictions of the thermoacoustic theory
239
A. Hofmann, S. Wild and L. R. Oellrich
Spontaneous oscillations of gas in a glass resonator: Observation of the local velocity and the simulation
243
A. Tominaga and T. Yazaki
Linear model of flow pattern for a valved three-stage pulse tube refrigeration
247
Z.-M. Xia, L.-M. Qiu, G.-B. Chen, L. Zhao, J.-Y. Zheng, J.-P. Yu and Z.-X. Huang
Radial temperature and velocity profiles of oscillating flows in a pulse tube refrigerator
251
K. Seo, M. Shiraishi, N. Nakamura and M. Murakami
Investigation of velocity profiles in oscillating flows inside a pulse tube refrigerator
255
M. Shiraishi, N. Nakamura, K. Seo and M. Murakami
Intrinsic behaviour of a four valve pulse tube refrigerator
259
M. Thfirk, H. Brehm, J. Gerster, G. Kaiser, R. Wagner and P. Seidel
Anomaly of one-stage double-inlet pulse tube refrigerator T. Shigi, Y. Fujii, M. Yamamoto, M. Nakamura, M. Yamaguchi, Y. Fujii, T. Nishitani, T. Araki, E. Kawaguchi and M. Yanai
263
Contents Temperature stability of pulse tube refrigerators
ix 267
N. Seki, S. Yamasaki, J. Yuyama, M. Kasuya, K. Arasawa, S. Furuya and H. Morimoto
Pulse tube refrigerator with low temperature switching valve
271
J.-T. Liang, C.-Q. Zhang, L. Xu, J.-H. Cai, E. Luo and Y. Zhou
On-off timing computer control system for valved refrigerator
275
L. Zhao, T. Sun, J.-Y. Zheng, Z.-X. Huang and G.-B. Chen
Increase in reservoir pressure of orifice pulse tube refrigerators
279
F. Kuriyama and Y. Fukasaku
Operation of a high-Tc SQUID gradiometer by use of a pulse tube refrigerator
283
G. Thummes, R. Landgraf, M. Mfick, K. Klundt and C. Heiden
Analysis and investigation of large diameter pulse tube refrigerator
287
Y.-Y. Guo, X.-X. Wang, S.-X. Bian and Y.-Z. Li
Active-buffer pulse-tube refrigerator
291
S.-W. Zhu, Y. Kakimi, K. Fujioka and Y. Matsubara
An inter-phasing pulse tube refrigerator for high refrigeration efficiency
295
J. L. Gao and Y. Matsubara
Experimental research on two-stage pulse tube refrigerator
299
T. Inoue, T. Matsui, S. Kawano and Y. Ohashi
Performance of the hybrid two-stage refrigerator
303
K. Tanida, J. L. Gao, Y. Hiresaki and Y. Matsubara
Improvement of the two stage pulse tube refrigerator reaching to 4 K
307
M. Tanaka, T. Nishitani, T. Kodama, T. Araki, E. Kawaguchi and M. Yanai
Nitrogen precooled multi-stage pulse tube refrigerator reaching 2.1 K
311
G. Thummes, S. Bender and C. Heiden
Thermodynamic calculation of three-stage pulse tube refrigerator
315
G.-B. Chen, Z.-M. Xia, L. Zhao, L.-M. Qiu, J.-Y. Zheng and J.-P. Yu
Three-staged pulse tube refrigerator with linear motor compressor
319
N. Yoshimura, Y. Matsubara, Y. Ohtani, H. Nakagome and H. Okuda
G - M Coolers
Effectiveness of magnetic regenerator material with low T~ below ~20 K regenerative character of magnetic materials
Very effective 325
T. Hashimoto
Development of a 4 K GM/JT refrigerator for Maglev vehicle
331
S. Fujimoto, S. Taneya, T. Kurihara, K. Miura, K. Tomioka, M. Okamoto, T. Yamaguchi, S. Kasahara, M. Terai, A. Miura, H. Nakao, T. Fujinami and Y. Nakamoto
Development of 2W class 4 K Gifford-McMahon cycle cryocooler
335
I. Takashi, N. Masashi, N. Kouki and Y. Hideto
Optimization of intake and exhaust valves for 4 K Gifford-McMahon cryocooler
339
R. Li, A. Onishi, T. Satoh and Y. Kanazawa
Investigation of the performance of a 4.2 K G-M refrigerator X.-D. Xu, L.-H. Gong, Z.-Y. Zhang and L.-A. Zhang
343
U.K.
Elsevier Science Ltd, The Boulevard, Langford Lane, Kidlington, Oxford, OX5 1GB, England
U.S.A.
Elsevier Science Inc., 660 White Plains Road, Tarrytown, New York, 10591-5153, U.S.A.
JAPAN Elsevier Science Japan, Tsunashima Building Annex, 3-20-12 Yushima, Bunkyo-ku, Tokyo 113, Japan
Copyright © 1997 Elsevier Science All Rights Reserved. No part of this publication may be reproduced, stored in a retrieval system or transmitted in any form or by any means; electronic, electrostatic, magnetic tape, mechanical photocopying, recording or otherwise, without prior permission in writingfrom the publisher. First edition 1997 Library of Congress Cataloging in Publication Data A catalogue record for this tit!e is available from the Library of Congress. British Library Cataloguing in Publication Data A catalogue record for this title is available from the British Library.
ISBN 0-08-042688-3
Printed and bound in Great Britain by BPC Wheatons Ltd, Exeter
Contents Cooling performance of a pressurized HeII cryogenics system for the superconducting magnet test facility at KEK N. Kimura, Y. Ajima, Y. Doi, T. Haruyama, N. Higashi, M. Iida, S. Kato, M. Kawai, H. Kawamata, S. Kim, Y. Kondou, Y. Makida, K. Mimori, S. Mizumaki, T. Nakamoto, T. Ogitsu, H. Ohhata, N. Ohuchi, S. Sugawara, T. Shintomi, K. Tanaka, A. Terashima, K. Tsuchiya, H. Yamaoka and A. Yamamoto
A supercritical/superfluid He-cryostat for STEP
xi
419
423
D. Mohr, A. Seidel, M. Sander, G. Jochimsen, A. Wagner, J. Wolf, J. Weber and K. Petersen
Design of thermal shield for the ITER cryostat K. Hamada, K. Nishida, T. Kato, T. Honda, H. Tsuji, A. Itoh, M. Nakahira, S. Shimamoto, M. E. P. Wykes, R. Bourque, I. Ohno and Y. Miyauchi
427
Design of superfluid-cooled cryostat for 1 GHz NMR spectrometer A. Sato, T. Kiyoshi, H. Wada, H. Maeda, S. Itoh and Y. Kawate
431
A variable temperature cryostat to measure Jnoncu(T) of ITER strands up to 20 teslas B. Jager, A. Bocquillon, P. Chaussonnet, A. Martinez, S. Nicollet, J. P. Serries and J. C. Vallier
435
Tests of the cryostat for 1.3 GHz superconducting cavity at T _< 1.8 K Yu. P. Filippov, A. M. Kovrizhnykh, V. I. Batin and S. V. Uchaikin
439
Comparison of floating and thermalized multilayer insulation systems at low boundary temperature G. Ferlin, B. Jenninger, P. Lebrun, G. Peon, G. Riddone and B. Szeless
443
Cooling Technique A LHE economiser at 1.8 K S. Buhler
449
Low noise gas flow cryosystem for cooling high-Tc squid J. Troell and C. Heiden
453
Compact dilution refrigerator S. Yoshida, S. Mori, T. Umeno, Y. Kamioka, M. Watanabe and Y. Ootuka
457
Experimental study of the dilution refrigerator without 1 K pot M. Maeda, T. Shigematsu, Z.-M. Li, T. Shigi, Y. Fujii, M. Yamaguchi and M. Nakamura
461
Numerical simulation of countercurrent heat exchangers in cryogenic systems M. Kauschke and H. Quack
465
The operation experience with the dual cooling system of the POLO coil (two phase-supercritical He) W. Herz, M. Siitier, A. Ulbricht and F. Wiichner
469
A vortex refrigerator for NMR experiments
473
M. Fujii, K. Nakamura, X. Xu, K. Kawano and K. Okada
Analytical model to calculate the transient thermo-mechanical behaviour of long thin structures cooled from a pipe: Application to the LHC dipole thermal shield G. Pe6n, G. Riddone and L. R. Williams Characterization of bonnetless cryogenic valves K. Daido, K. Yoshikawa, R. Maekawa and S. Satoh Flexible corrugated cryotransferlines, long term experience at JET and the experience with supercritical helium flow conditions W. Obert and C. Mayaux
477
481
485
xii
Contents
Experimental results with superinsulated cryogenic transfer line test modules in THISTA IV. Lehmann, M. Seidler and M. Stamm
489
Design and construction of long cryogenic piping lines K. Kawano, K. Hamada, T. Kato, T. Honda, K. Nishida, K. Matsui, T. Hiyama, K. Ohtsu, S. Sekiguchi, H. Tsuji, M. Ando, T. Hiyama and K. Ichige
493
Calculating radiation exchange factors of radiant cryocooler by the Monte Carlo method H. Yao, C.-H. Gao, C.-Z. Chen and Y.-Z. Li
497
Development of loop heat pipes for cryogenic applications R. Chandratilleke, Y. Ohtani, H. Hatakeyama and H. Nakagome
501
New cryosurgical probe suitable for endoscopic application in the minimal invasive therapy B. Schumann, A. Binneberg and R. Herzog
505
Thermosiphon cooler: A low microphonic cooling system for HTC-devices; especially for SQUIDs A. Binneberg, H. Buschmann, R. Herzog, J. Neubert and G. Sp6"rl
509
What happened to cryogenic and superconducting equipment in the Great Hanshin Earthquake? Final report A. Sato, K. Fujioka, T. Haruyama, H. Hirabayashi, T. Ishigohka, A. Ishiyama, Y. Kawate, K. Nishigaki, S. Nishijima, T. Noguchi, O. Ogino, H. Ogiwara and O. Okazaki
513
Superfluid Helium Thermohydraulic behaviour of HelI in stratified co-current two-phase flow B. Rousset, A. Gauthier, L. Grimaud, A. Bezaguet and R. van Weelderen
519
Influence of fountain pressure on heat transfer to superfluid helium Y.-Z. Li, U. Ruppert, L Arend and K. Lfiders
523
Analysis and characterization of saturated bath He II heat exchangers S. W. Van Sciver and S. J. Welton
527
Experimental study of propagation characteristics of He i-He II interface K. Kamiya, M. Murakami and N. Yanagise
531
Steady and unsteady heat transfer from a horizontal wire with a wide range of diameter in a pool of subcooled He II at pressures M. Shiotsu, K. Hata, Y. Takeuchi, K. Hama and A. Sakurai
535
Two dimensional heat transport in He II channel including a copper wall T. Okamura, S. Hamaguchi, S. Sakuma, T. Suekane and S. Kabashima
539
Numerical investigation of evolution of vortex line density in the case of transient heating T. Kanari and M. Murakami
543
Heat transfer from superconductor wire to superfluid helium Y.-Z. Li, Y.-Z. Wu, Y.-Y. Wu, U. Ruppert, I. Arend and K. Lfiders
547
The effect of spacer arrangement on the heat transfer in He I and He II channels H. Kobayashi and K. Kawakami
551
Pressure effect on the heat transfer in bath of superfluid helium R.-Z. Wang, P. Zhang and J.-Y. Wu
555
Kapitza conductance of niobium for S. R. F. cavities A. Boucheffa, M. X. Franqois and J. Amrit
559
Contents
xiii
Thermal behaviour of electrical multilayer insulation permeable to superfluid helium B. Baudouy, A. Boucheffa, M. X. Fran,cois and C. Meuris
563
Pressure gradient caused by quantized vortex in superfluid helium M. Yamaguchi, Y. Fujii, M. Nakamura, T. Shigematsu and T. Shigi
567
Heat and mass transfer between two saturated He II baths X. Huang, J. Panek and S. W. Van Sciver
571
Measurement of characteristic time of quantized vortex development using a thermal shock wave T. Shimazaki and M. Murakami
575
Heat Transfer Non-dimensional correlation for boiling heat transfer from sintered porous layer surface R.-S. Wang, A.-Z. Gu, Z. Li and J.-H. Huang
581
Critical heat fluxes in pool boiling of subcooled liquid nitrogen at elevated pressures K. Hata, M. Shiotsu and A. Sakurai
585
The measurement of vapor bubble vibration during noisy film boiling in superfluid helium M. Yamaguchi and M. Murakami
589
Influence of surface roughness on transient nucleate boiling of cryogens S. Fuchino, N. Tamada, I. Ishii and M. Okano
593
Superheating of liquid mixtures of 3He and 4He K. Nishigaki, M. Takeda and Y. Maruno
597
Nucleate pool boiling heat transfer to slush hydrogen K. Ohira and H. Furumoto
601
Heat transfer characteristics of a prototype pool boiling superconductor to liquid helium A. Iwamoto, T. Mito, K. Takahata, N. Yanagi and J. Yamamoto
605
Temporal deterioration of helium heat transfer at moderate pulse heat load Y. P. Filippov and I. A. Sergeyev
609
Cryogenic characteristic investigation on heat transfer between gas and solids of an adiabatic moving-bed Y.-Y. Guo, Z.-Z. Li, L. Wang and L.-A. Zhang
613
Transient heat transfer from a silver sheathed high-Tc superconducting tape in liquid nitrogen M. Shiotsu, K. Hata, A. Sakurai, C. Suzawa, S. Isojima and K. Sato
617
Surface treatment of aluminum heat switch T. Shigematsu, M. Maeda, M. Takeshita, Y. Fujii, M. Nakamura, M. Yamaguchi, T. Shigi and H. Ishii
621
The study on the solid thermal contact resistance at low temperatures L. Xu, S.-L. Zhou, J. Yang and J.-M. Xu
625
Experimental study on thermal contact conductance at liquid helium temperature K. Sunada and Y.-M. Kang
629
Gas Properties On the Joule-Thomson integral inversion curves of helium-3, helium-4 and hydrogen B.-Z. Maytal and A. Shavit
635
xiv
Contents
A modified Patel-Teja equation of state for cryogenic fluids G.-M. Chen, Z.-Z. Yin and G.-B. Chen
639
Helium extraction from thermal spring gases D. Ghose, B. Sinha, R. Dey, S. K. Das and D. G. Bhattacharya
643
Measurements Temperature measurement under high magnetic fields around 1.8 K by using CGR thermometers T. Haruyama, N. Kimura, K. Tanaka and A. Yamamoto
649
New type of thin-film germanium resistance thermometer for use in a wide temperature range V. Mitin, Y. Tkhorik and E. Venger
653
Thermoacoustic Taconis oscillations in helium-4 liquid level sensors S. Yoshida, K. V. Ravikumar, J. Williamson, N. Papavasiliou and T. H. K. Frederking
657
Use of strain gages for low temperature thermal expansion measurements R. P. Walsh
661
Strain measurements of stainless steel at low temperatures using electronic speckle pattern interferometry S. Nakahara, H. Sakiyama, S. Hisada and T. Fujita
665
Friction and wear testing at cryogenic temperatures T. Gradt, W. Hfibner and H. B6rner
669
A flexible and extensible monitoring system for large scale experiments J. Kariya, H. Okumura, M. Emoto, M. Shoji, Y. Teramachi, T. Ohska, Y. Shamoto, S. Yamaguchi, O. Motojima and J. Yamamoto
673
Development of a low temperature measurement and control system for measuring liquid oxygen density D. Wang, Z.-X. Huang, G.-B. Chen, J.-Y. Zheng, J.-P. Yu, Y.-Y. Li, M. Shen, H.-Y. Chen, R.-L. Xu, G.-W. Cui and R.-M. Liu
677
Investigation of thermal and vacuum transients on the LHC prototype magnet string P. Cruikshank, N. Kos, G. Riddone and L. Tavian
681
Innovative device producing double-layer cryogenic pellet H. Itoh and S. Sudo
685
A new concept to detect He leaks on TORE SUPRA cryogenic plant F. Minot, G. Bon Mardion, B. Jager, B. Gravil, J. L. Marechal and J. L. Violet
689
One consideration for recovery heat flux of directly heated wires Y.-Z. Li, Y.-Z. Wu, Y.-Y. Wu, U. Ruppert, L Arend and K. Lfiders
693
A novel superconductor V-I simulator for 77 K B. ten Haken and H. H. J. ten Kate
697
The estimation of critical current density using SRPM and AC methods S. Koto, H. Nakane, E. S. Otabe, T. Matsushita, S. Nagaya and S. Yoshizawa
701
Contents
xv
PART TWO Space Cryogenics ISO
707
ISO in-orbit cryogenic performances, system aspects
711
Half a year on orbit Th. Paflvogel, A. Seidel, U. Sagner, G. Jochimsen, M. Sander, J. Wolf, E. Ettlinger and K Petersen B. Collaudin, Th. Passvogel, A. Seidel and J. J. Juillet
The FIRST-cryosystem with ISO-technology Th. Paflvogel, A. Seidel, J. Schupp, M. Sander and G. Jochimsen
715
Study of cryodeposition contamination of radiant coolers D. Dong and W.-Y. Wang
719
Application of Superconductivity Fusion Large helical device project for SC steady-state fusion experiment O. Motojima
725
Superconducting coil system for the LHD and its R&D J. Yamamoto and LHD Group
731
Cooling and excitation experiments of a single inner vertical coil (EXSIV) for the large helical device T. Satow, T. Mito, K. Takahata, J. Yamamoto, S. Satoh, A. Nishimura, S. Tanahashi, S. Yamada, H. Chikaraishi, H. Tamura, N. Yanagi, S. Imagawa, A. Iwamoto, R. Maekawa, K. Yamazaki, S. Yamaguchi, K. Watanabe, N. Inoue, H. Suzuki, T. Morisaki, S. Masuzaki, Y. Katoh, S. Sakakibara, K. Nishimura, M. lima, H. Hayasi, T. Baba, S. Moriuchi, K. Ohba, H. Ogawa, H. Sekiguchi, I. Ohtake, O. Motojima, M. Takeo, A. Ninomiya, S. Ioka and T. Uede Hydraulic characteristics and stability for the experiments on a single inner vertical coil (EXSIV) for the large helical device K. Takahata, T. Mito, T. Satow, A. Nishimura, S. Yamada, H. Chikaraishi, N. Yanagi, A. Iwamoto, R. Maekawa, S. Imagawa, H. Tamura, S. Satoh, S. Tanahashi, J. Yamamoto, O. Motojima, EXSIV Group, T. Kai, K. Nakamoto, M. Ono and T. Yoshida
735
739
Test facilities of the experiments on a single inner vertical coil (EXSIV) for the large helical device T. Mito, K. Takahata, T. Satow, S. Satoh, A. Nishimura, S. Yamada, N. Yanagi, H. Chikaraishi, A. Iwamoto, R. Maekawa, S. Tanahashi, T. Kai, T. Yamamoto, Y. Wachi, T. Uede, H. Hiue, I. Itoh, N. Saji, A. Miyake, I. Ohno, O. Motojima, J. Yamamoto and EXSIV Group
743
Electrical power system for EXSIV H. Chikaraishi, S. Yamada, T. Mito, T. Satow, S. Tanahasi, O. Motojima, J. Yamamoto, EXSIV Group, T. Uede, H. Hiue, Y. Yonenaga and S. Araki
747
Development, fabrication, testing and joints of aluminum stabilized superconductors for the helical coils of LHD N. Yanagi, T. Mito, S. Imagawa, T. Satow, K. Takahata, A. Nishimura, S. Yamada, A. Iwamoto, S. Yamaguchi, H. Chikaraishi, J. Yamamoto, O. Motojima, LHD Group, Y. Kuchiishi, T. Tamaki, T. Senba, K. Asano, S. Suzuki, T. Miyaji, S. Inaba, M. Seido and H. Moriai Deformation analysis for coil pack simulating large scale pool-boiling superconducting coil H. Tamura, A. Nishimura, S. Imagawa, J. Yamamoto, K. Asano, T. Tamaki, K. Nakanishi and S. Suzuki
751
755
xvi
Contents
Deformation behavior of coil pack for helical coil in large helical device A. Nishimura, H. Tamura, S. Imagawa, J. Yamamoto, K. Asano, T. Tamaki, K. Nakanishi and S. Suzuki
759
ITER CS model coil project N. Mitchell, K. Okuno, R. Thome, H. Tsuji, T. Ando, S. Shimamoto, J. Jayakumar and J. Minervini
763
Design of the ITER-CS model coil ~
767
Joint development of the CS insert coil M. Sugimoto, A. Terasawa, H. Nakajima, T. Kato, Y. Nunoya, K. Matsui, Y. Takahashi, T. Ando, T. Isono, N. Koizumi, H. Tsuji, S. Shimamoto, T. Ichihara, T. Sasaki, M. Hasegawa and T. Minato
771
Development of 46 kA layer to layer joint for INTER-CS model coil Y. Nunoya, Y. Takahashi, T. Ando, H. Nakajima, T. Kato, M. Sugimoto, M. Oshikiri, T. Isono, N. Koizumi, K. Matsui, Y. Miura, H. Tsuji, S. Shimamoto, H. Ogata and J. Shibuya
775
AC loss measurement of 46-kA conductor joint for CS model coil K. Matsui, T. Ito, Y. Takahashi, Y. Nunoya, M. Sugimoto, H. Fujisaki, H. Tsuji, and S. Shimamoto
779
Fabrication R & D of outer module of ITER CS model coil T. Sasaki, S. Yanaka, M. Ushijima, S. Murai, T. Hirumachi, J. Inagaki, H. Ogata, Y. Sumiyoshi, O. Ohsaki, T. Fujioka, T. Ando, Y. Takahashi, H. Nakajima, M. Sugimoto, K. Hamada and H. Tsuji
783
Consideration of current sharing temperature measurement for ITER CS model coil T. Honda, T. Ando, H. Tsuji and S. Shimamoto
787
Cool-down simulation of 46 kA and 13 T Nb3A1 insert N. Koizumi, T. Ito, T. Ando, M. Sugimoto, A. Terasawa, M. Nozawa, L Watanabe, H. Tsuji, S. Shimamoto, K. Okuno, H. Tsukamoto and M. Otsuka
791
Experimental results and design of 50-kA forced-cooled current leads for fusion machines Y. Takahashi, M. Sugimoto, K. Matsui, K. Takano, M. Nozawa, Y. Nunoya, N. Koizumi, H. Nakajima, T. Kato, K. Nishida, T. Honda, T. Ando, H. Tsuji, Y. Yasukawa, L Itoh, K. Sakaki and M. Konno
795
RF jacketing line for manufacturing ITER cable-in-conduit conductor
799
outer module T. Ando, Y. Takahashi, H. Nakajima, T. Kato, T. Isono, M. Sugimoto, K. Hamada, N. Koizumi, Y. Nunoya, K. Matsui, A. Terasawa, K. Nishida, T. Honda, H. Tsuji, S. Shimamoto, T. Fujioka and O. Osaki
V. Sytnikov, I. Peshkov, A Taran, I. Semeonov, A. Rychagov, V. Mitrohin, P. Dolgosheev and L Chensky Fast ramp 50 kA superconducting transformer for testing full-size ITER cable joints
803
H. G. Knoopers, S. Wessel, H. J. G. Krooshoop, O. A. Shevchenko, A. Godeke, H. H. J. ten Kate, B. A. Smith, R. J. Camille and J. V. Minervini Plasma disruption simulation test facility
807
T. Ando, S. Iwamoto, I. Itou, H. Tsuji and S. Shimamoto Cool down and charging tests of a double-conduit type forced-flow superconducting coil H. Morita, Y. Wadayama, Y. Hotta, K. Kouriki, Y. Murata, R. Takahashi, K. Asano, A. Shigenaka, Y. Takahashi, K. Yoshida and H. Tsuji
811
Cooling of the W7-X superconducting coils F. Schauer
815
Contents Test facility in support of the SST-1 CICC
xvii 819
B. Sarkar, S. Pradhan, C. P. Dhard, V. Tanna, A. K. Sahu, A. Amardas, Y. C. Saxena and S S T Team
Analysis of CICC behaviour in the SST-1 magnets
823
S. Pradhan, S. Das, B. Biswas, Y. C. Saxena and S S T Team
Energy dissipation in SST-1 coil joints
827
S. Pradhan, S. Das, B. Biswas, Y. C. Saxena and S S T Team
Stabilization and confinement of discharge plasma with high-Tc bulk superconductor tubes (Supertrons) H. Matsuzawa, Y. Mizutani, J. Ishikawa and S. Suganomata
831
Accelerators and Detectors
Power test results of the first LHC second generation superconducting single aperture lm long dipole models
837
A. Siemko, G. Kirby, J. Ostler, D. Perini, N. Siegel, D. Tommasini and L. Walckiers
Testing of TRISTAN insertion quadrupole magnet in superfluid helium
843
K. Tsuchiya, T. Ogitsu, N. Ohuchi, K. Sasaki, Y. Doi, T. Haruyama, S. Kato, H. Kawamata, S. Kim, N. Kimura, N. C. Bhattacharya, T. Shintomi, K. Tanaka, A. Terashima and A. Yamamoto
A superconducting wiggler magnet for the National Laboratory of Synchrotron Radiation
847
Q. L. Wang, Y. F. Bi, B. Z. Li, L. R. Din, F. T. Wang, C. W. An and L. Z. Lin
Fabrication and testing of a 13 m-long SSC model dipole magnet K. Hosoyama, K. Hara, N. Higashi, A. Kabe, H. Kawamata, Y. Kojima, Y. Morita, H. Nakai, A. Terashima, T. Takahashi, T. Shintomi, H. Hirabayashi and Y. Kimura
851
Design study of the injection system of the RIKEN superconducting ring cyclotron
855
H. Okuno, T. Tominaka, T. Kubo, J.-W. Kim, T. Mitsumoto, T. Kawaguchi, Y. Tanaka, S. Fujishima, K. Ikegami, A. Goto and Y. Yano
Design study of sector magnet for RIKEN superconducting ring cyclotron
859
T. Kawaguchi, T. Kubo, T. Mitsumoto, T. Tominaka, S. Fujishima, H. Okuno, Y. Tanaka, J.-W. Kim, K. Ikegami, A. Goto and Y. Yano
The aluminum stabilized conductor for the Fermilab DO solenoid
863
R. P. Smith, H. E. Fisk, K. Krempetz, R. Yamada, S. Mine, T. Kobayashi, L L. Horvath, H. P. Marti, J. Neuenschwander, D. Grman, R. Huwiler, H. Eriksson, J. Seppala, J. Teuho, W. B. Sampson, A. K. Ghosh, B. Seeber, L. Erbuke and R. Flukiger
Ten year operational experience of the TRISTAN detector selenoid magnets
867
T. Haruyama, Y. Doi, Y. Makida, K. Aoki, T. Kobayashi, Y. Kondo, M. Kawai, K. Tsuchiya, M. Wake and A. Yamamoto
Cryogenic system for the Muon g-2 Superconducting Magnet
871
X.-J. Lin, G. Bunce, J. R. Cullen, M. A. Green, C.-I. Pai, L. P. Snydstrup and T. Tallerico
Forced two-phase helium cooling of four 15-meter-in-diameter superconducting solenoids
875
X.-J. Lin, G. Bunce, J. R. Cullen, M. A. Green, C.-I. Pai, L. P. Snydstrup and T. Tallerico
Gas cooled cryogenic power leads for the relativistic heavy ion collider M. L. F. Rehak and A. Nicoletti
879
xviii
Contents
Motors and Generators Present status and future prospect of research & development of superconducting technology for electric power apparatuses- mainly on R&D for superconducting generator S. Nakayama, M. Kazumori and K. Komatsu
885
Applications of superconducting technology to electric power systems Y. Aiyama, S. Nakayama, M. Nishikawa, T. Ageta and T. Nitta
891
Investigation of superconducting turbogenerator operation in the network L. I. Chubraeva
895
R&D project on superconducting generators in Super-GM T. Ageta
899
Development of 70MW class superconducting generators K. Yamaguchi, K. Suzuki, K. Miyaike, K. Toyoda and T. Ichikawa
905
A series of studies on superconducting generator with high response excitation T. Kisida, Y. Imai, T. Nitta, T. Okada, H. Hasegawa and H. Nagamura
909
Investigation of superconductor for 70 MW class model superconducting generator of quick response type H. Takigami, H. Nakamura, E. S. Yoneda, M. Sugimoto, A. Kimura, H. Sakamoto, Y. Furuto, K. Inoue and K. Sato Analysis of AC losses for superconductors for 70 MW class model superconducting generator of quick response type A. Kimura, M. Sugimoto, H. Sakamoto, Y. Furuto, H. Takigami, H. Nakamura, E. S. Yoneda, K. Inoue and K. Sato
913
917
AC losses in a double-stranded Nb-Ti cable for a quick-response-type superconducting generator M. Sugimoto, A. Kimura, H. Sakamoto, Y. Furuto, H. Takigami, H. Nakamura, E. S. Yoneda, F. Sumiyoshi, S. Kawabata, K. Inoue and K. Sato
921
Experimental investigation of helical armature winding for a light cryoalternator of 1-3 MW class L. I. Chubraeva and D. V. Sirotko
925
Combination of HTSC, high-purity aluminum and Nd-Fe-B in a synchronous alternator L. I. Chubraeva, I. S. Ganzinov, S. N. Pylinina, V. A. Sapozhnicov, V. E. Sigaev and V. A. Tutaev
929
Critical current and strain of in-situ processed Nb3Sn superconductor under cyclic mechanical load K. Goto, S. Iwaski, N. Sadakata, T. Saitoh, O. Kohno, S. Torii, H. Kasahara, S. Akita and J. Yoshitomi
933
Design and construction of an HTSC synchronous machine with permanent magnet excitation L Vajda, A. G. Mamalis and A. Szalay
937
Machine constants of 30 kVA superconducting synchronous motor with rotating field windings K.-D. Choi, T. Hoshino, H. Tsukiji, M. Tsukiyama, T. Nishiya, I. Muta, E. Mukai and S.- Y. Hahn
941
HTSC bulk magnet motor Z. Szfics and U. Ruppert
945
Rotating helium transfer coupling for a 200 KVA superconducting generator S. Jacob, S. Kasthurirengan, R. Karunanithi, T. Suryanarayana, K. A. Durga Prasad, K. S. N. Raju and J. L. Bhattacharya
949
Contents
xix
Test of high-voltage output magnetic flux pump for superconducting magnet H. Tsukiji, K.-D. Choi, M. Tsukiyama, T. Nishiya, T. Hoshino, I. Muta and E. Mukai
953
Dielectric breakdown of liquid helium in a rotating cryostat L Ishii, S. Fuchino, M. Okano and N. Tamada
957
Power Cables Research and development of compact high-Tc superconducting cables T. Hara, H. Ishii and S. Honjo
963
The development of HTS cable technology T. Shibata, M. Watanabe, C. Suzawa, S. Isojima, J. Fujikami, N. Saga, K. Ohmatsu, K. Sato, H. Ishii, S. Honjyo and T. Hara
967
Development of the termination for the 77 kV-class high Tc superconducting power cable T. Masuda, S. Isojima, T. Shimonosono and S. Nagaya
971
Development of long length HTSC cable conductor J. Fujikami, N. Saga, K. Ohmatsu, T. Shibata, M. Watanabe, C. Suzawa, S. Isojima, K. Sato, H. Ishii, S. Honjo and T. Hara
975
Research and development of 50 m-long high-Tc superconductor for power cables S. Mukoyama, K. Miyoshi, H. Tsubouti, M. Mimura, N. Uno, Y. Tanaka, N. Ichiyanagi, H. Ishii, S. Honjo and T. Hara
979
1.5 kA-class multilayered Bi-2223/Ag conductors: DC and AC characterizations M. Luciano, C. Franco, M. Ernesto and O. Vanni
983
Fabrication and evaluation of the high-Tc conductors using Bi-2223 Ag sheathed tapes A. Kume, S. Nagaya, T. Shimonosono, M. Nakagawa, M. Nagata, N. Sadakata, T. Saitoh, O. Kohno and M. Ono
987
R&D on oxide superconducting wires for application to power apparatuses N. Yoshida, K. Sato, J. Yoshitomi and T. Ichikawa
991
Fabrication of YBCO superconducting tape by continuous MOCVD technique K. Onabe, S. Nagaya, T. Shimonosono, Y. Iijima, N. Sadakata, T. Saito and O. Kohno
995
Alignment of unidirectionally grown YBCO S. Asakura, S. Nagaya, T. Shimonosono, M. Nakagawa, N. Sadakata, T. Saitoh and O. Kohno
999
Fabrication of high Jc YBCO tapes using YSZ buffer layers deposited by IBAD method M. Hosaka, Y. Iijima, N. Sadakata, T. Saito, O. Kohno and J. Yoshitomi
1003
Power Devices Design and construction of a 500 kVA-class oxide superconducting power transformer cooled by liquid nitrogen K. Funaki, M. Iwakuma, M. Takeo, K. Yamafuji, J. Suehiro, M. Hara, M. Konno, Y. Kasagawa, K. Okubo, Y. Yasukawa, S. Nose, M. Ueyama, K. Hayashi and K. Sato
1009
Characteristics and AC losses of coreless superconducting autotransformers K. Kajikawa, K. Kaiho, M. Yamamoto, H. Fuji, N. Sadakata, T. Saito and O. Kohno
1013
Design study for the development of a 1 kA persistent-current-switch in a type of transformer H. Hayashi, T. Imayoshi, K. Tsutsumi, M. Takeo, S. Sato, H. Morimoto, K. Asano and K. Yamaguchi
1017
xx
Contents
Inductive superconducting fault curren[ limiter development J. R. Cave, D. W. A. Will, n, R. Nadi, W. Zhu and Y. Brissette
1021
Magnetic shield type high Tc superconducting fault current limiter T. Onishi, S. Yamasaki and A. Nii
1025
Development of 6.6 kV/1 kA single-phase superconducting fault current limiter loss reduction T. Yazawa, T. Kurusu, M. Takahashi, K. Yamamoto, S. Nomura, M. Urata, T. Ohkuma, M. Nakade and T. Hara
AC 1029
Development of a 6.6 kV class compact fault current limiter M. Takahashi, H. Nakagome, M. Urata, T. Ohkuma, M. Nakade and T. Hara
1033
Electrical application of a HTS saturable magnetic core fault current limiter J. X. Jin, C. Grantham, X. Y. Li, Z. Y. Liu, Y. C. Guo, T. R. Blackburn, H. L. Liu, J. N. Li, H. K. Liu, Z. J. Zeng, J. Y. Liu and S. X. Dou
1037
Progress of 480 MJ/40 MW SMES component development project Y. Murakami, O. Tsukamoto, T. Sato, E. Masada, K. Ogiso, S. Neo and M. Hosokawa
1041
A conceptual design of a concentric double spherical coil system for a medium scale SMES T. Ezaki, N. Watanabe and I. Kamiya
1045
Design study for the development of 1 kWh/1 MW module type SMES T. Imayoshi, K. Tsutsumi, F. Irie, M. Takeo, K. Funaki, H. Okada, T. Ezaki, R. Itoh, F. Sumiyoshi, Y. Satoh, K. Asano and S. Nose
1049
Experimental results of cryogenic stability of the superconductor for SMES A. Tomioka, T. Bohno, S. Nose, M. Konno, K. Sakaki, M. Takeo, K. Funaki, S. Sato and M. Matsuo
1053
Fundamental study of a HTS coil for SMES T. Masuda, S. Isojima, K. Ohkura, K. Sato, A. Ryouman, T. Kaito, T. Kishida and S. Uno
1057
Advanced mechanical switch for persistent current of SMES S. Nozaki, M. Masuda and M. Hamada
1061
Design and considerations on superconducting shunt reactors T. Nitta, S. Nogawa, Y. Nisiwaki and H. Nomura
1065
A proposal of new excitation method of superconducting magnet using LC resonance T. Ishigohka, K. Shimizu, A. Ninomiya and T. Koga
1069
Test results of a 200 A class fast response magnetically controlled persistent current switch H. Kimura, K. Noto, M. Matsukawa, S. Fujinuma, T. Segawa, T. Takahashi, N. Sadakata, T. Saito, K. Goto, O. Kohno, H. Honma and C. Takahashi
1073
Magnetic Levitation Design study of a superconducting bulk magnet for attractive-type magnetic levitation H. Ohsaki, A. Senba and E. Masada
1079
Magnetically levitated transport system in vacuum using high-Tc superconductors H. Minami, N. Ueda, T. Koike and J. Yuyama
1083
High Field Magnets High field magnets of Tsukuba Magnet Laboratories K. Inoue, T. Kiyoshi, T. Asano, Y. Sakai, G. Kido, H. Wada and H. Maeda
1089
Contents xxi Development of 18 T high magnetic field superconducting magnet O. Ozaki, M. Yoshikawa, R. Hirose, T. Miyazaki, T. Miyatake, M. Shimada, K. Matsumoto, M. Hamada and K. Takabatake
1095
N R I M R&D program on HTS coils for 1 GHz N M R spectrometer
1099
T. Kiyoshi, K. Inoue, M. Kosuge, H. Kitaguchi, H. Kumakura, H. Wada and H. Maeda
Development project of 1 GHz N M R spectrometer K. Inoue, T. Kiyoshi, A. Sato, K. Itoh, H. Wada, H. Maeda, R. Ogawa, Y. Kawate, K. Takabatake, T. Horiuchi, J. Kida and K. Higuchi
1103
Cryocooler-cooled Magnets A cryocooler cooled 10.7 T superconducting magnet with a room temperature bore of 52mm K. Jikihara, K. Watazawa, J. Sakuraba, T. Hasebe, H. Mitsubori, M. Ishihara, Y. Yamada, K. Watanabe and S. A waji
1109
Development of a 11.5 T liquid helium-free superconducting magnet system Y. Ohtani, H. Hatakeyama, H. Nakagome, K. Koyanagi, T. Yazawa and S. Nomura
1113
A cryofree superconducting magnet for industrial applications K. Timms, P. Daniels, M. Wade and J. Boehm
1117
Cryocooler-cooled large bore NbTi superconducting magnet using high temperature superconducting current leads K. Watanabe, T. Masumoto, S. Awaji, L Mogi, M. Motokawa, K. Watazawa, J. Sakuraba, T. Hasebe, M. Ishihara and Y. Yamada
1121
A superconducting magnetic separator with an integral refrigerator for blue-green algae H. Isogami, N. Saho and M. Morita
1125
Design and fabrication of cryogen free superconducting magnet K. Shibutani, S. Itoh, T. Takagi, O. Ozaki, R. Horise, S. Hayashi, M. Shimada, R. Ogawa, Y. Kawate, Y. Inoue, K. Matsumoto and N. Kimura
1129
Persistent current switch for a conduction cooled superconducting magnet T. H. Kim, S. Yokoyama and S. Yamamoto
1133
Current Leads Potential of high-temperature superconductor current leads for LHC cryogenics A. Ballarino, A. Ijspeert and U. Wagner Design of 12.5 kA current leads for the Large Hadron Collider using high temperature superconductor material A. Ballarino, A. Ijspeert, M. Teng, U. Wagner, S. Harrison, K. Smith and L. Cowey
1139
1143
Design and tests on the 30 to 600 A HTS current leads for the Large Hadron Collider A. Ballarino and A. Ijspeert
1147
Development of HTS current leads for 1 kWh/1 MW module type SMES system T. Bohno, A. Tomioka, S. Nose, M. Konno, K. Sakaki, T. Uede, T. Imayoshi, H. Hayashi, K. Tsutsumi and F. Irie
1151
Development of kA-class superconducting assembly conductor and current lead O. Kasuu, K. Takahashi, K. Sato and N. Yoshida
1155
A proposal for a Peltier current lead
1159
S. Yamaguchi, K. Takita and O. Motojima
xxii
Contents
Transient analysis of high-temperaturesuperconducting current leads T. Nishioka, K. Maehata, K. Ishibashi, M. Takeo, T. Mito and J. Yamamoto
1163
SQUID and Electronic Devices A plastic molded DC-SQUID for biomagnetic measurement Y. Utaka and T. Kido
1169
Development of SQUID based systems cooled by GM/JT cryocoolers K. Sata, S. Fujimoto, N. Fukui, E. Haraguchi, T. Kido, K. Nishiguchi and Y.-M. Kang
1173
Ferromagnetic fluxgate for measurement of the weak magnetic field structure with the use of HTSC - SQUID S. I. Bondarenko
1177
Multichannel high-Tc SQUID based heart scanner H. J. M. ter Brake, J. Flokstra, D. Veldhuis and H. Rogalla
1181
High current density NbN/A1N/NbN tunnel junctions as submillimeter wave SIS mixers Z. Wang, Y. Uzawa and A. Kawakami
1185
Fabrication of SrTiO3/YBa2Cu307_x heterostructure by ion beam sputtering T. Saito, X. Cai, K. Usami, T. Kobayashi and T. Goto
1189
Normal current controllability of JOFETs using InAs-inserted-channel InA1As/InGaAs inverted HEMTs T. Akazaki, H. Takayanagi, J. Nitta and T. Enoki
1193
Andreev reflection spectroscopy of superconducting mesoscopic devices Y. Misaki, A. Saito, K. Goto, M. Mikawa and K. Hamasaki
1197
Infrared radiation detector using YBCO thin film Y. Kakehi, T. Yotsuya, T. Kusaka, Y. Suzuki, S. Ogawa and H. Imokawa
1201
YBCO thin film terahertz radiator excited by femtosecond laser pulse M. Hangyo, S. Tomozawa, Y. Murakami, M. Tonouchi, M. Tani, Z. Wang and K. Sakai
1205
YBCO thin film bow-tie antenna for terahertz radiator
1209
N. Wada, M. Hangyo, Y. Murakami, M. Tonouchi, M. Tani, Z. Wang and K. Sakai
Stability and AC Loss Current redistribution in multistrand Nb3Sn CICC due to current ramp V. Vysotsky, M. Takayasu, S.-K. Jeong, P. Michael, J. Schultzand and J. Minervini Direct measurements of current distribution in a 12 strand Nb3Sn CICC (Part 1, experimental set-up) V. Vysotsky, M. Takayasu, S.-K. Jeong, P. Michael, J. Schultz and J. Minervini Direct measurements of current distribution in a 12 strand Nb3Sn CICC (Part 2, experimental results) V. Vysotsky, M. Takayasu, S.-K. Jeong, P. Michael, J. Schultz and J. Minervini
1215
1219
1223
Effects of current distribution on the stability of a triplet NbTi/Cu superconductor N. Hirano, T. Mito, K. Takahata, A. Iwamoto, R. Maekawa and J. Yamamoto
1227
Voltage spikes in ramped field experiments S.-K. Jeong, V. Vysotsky, M. Takayasu, J. H. Sehultz, P. C. Michael, S. Shen and W. Warnes
1231
Contents
xxiii
Contact resistance measurement of superconducting strands T. S. Jaffery, M. Wake, R. Scanlan and P. Mclntyre
1235
Measurements of cross contact resistance in Rutherford cables T. Suzuki, T. Shioiri, A. Ishiyama and K. Hosoyama
1239
Effect of Cr plating on the coupling current loss in cable-in-conduit conductors P. Bruzzone, A. Nijhuis and H. H. J. ten Kate
1243
AC loss measurement of superconducting dipole magnets by the electrical method Y. Morita, K. Hara, N. Higashi, K. Hosoyama, A. Kabe, H. Kawamata, Y. Kojima, H. Nakai, T. Shintomi and A. Terashima
1249
Measurement of time constants for coupling losses in the LHD superconductors N. Yanagi, S. Tak6cs, T. Mito, K. Takahata, A. Iwamoto and J. Yamamoto
1253
Coupling losses, time constants and current distribution in superconducting cables in spatially changing magnetic field S. Takfcs AC losses measurement in cable-in-conduit conductor by calorimetric method Y. Wadayama, H. Morita, K. A ihara and R. Takahashi
1257
1261
Void fraction effect on stability in cable-in-conduit conductor for varying field due to plasma disruption A. Terasawa, Y. Miura, T. Ando, M. Nozawa, T. Isono, Y. Takaya, H. Tsuji and S. Simamoto
1265
The influence of Cu/SC ratio on the stability of superconducting magnet impregnated by epoxy-resin Q.-L. Wang, C.-W. An, Y.-Q. Wang and L.-Z. Li
1269
Relaxation of superconducting wire tension and magnet stability H. Moriyama, F. Sawa, H. Mitsui, M. Arata, S. Nishijima and T. Okada
1273
Calculation of wire motion in a superconducting magnet T. Kushida, S. Nishijima, Y. Honda and T. Okada
1277
Numerical experiments for transient stability in dry-winding superconducting magnets S.-B. Kim and A. Ishiyama
1281
One dimensional simulation of normal state propagation in a superconducting cable M. Emoto, T. Senba, S. Yamaguchi, N. Yanagi and O. Motojima
1285
Stability of synthetic enameled superconducting wires K. Seo, M. Morita and S. Yamamoto
1289
Quench behavior in superconducting solenoid magnets R. Hirose, T. Kamikado, 0. Ozaki, M. Yoshikawa, M. Hamada and K. Takabatake
1293
A study on stability of conductors under a high gravitational field N. Higuchi, K. Kaiho, I. Ishii, H. Nomura, H. Tateishi, S. Fuchino, K. Arai, S. Sekine, N. Natori, K. Kajikawa, N. Tamada, K. Tsugawa and K. Fujima
1297
Evaluation of inductive heating energy of quench experiment on a long length (QUELL) conductor with the calorimetric method T. Ito, N. Koizumi, Y. Miura, K. Matsui, H. Wakabayashi, Y. Takahashi, H. Tsuji and S. Shimamoto Distribution of superconducting currents in NbTi/Nb/Cu multilayered cylinder S. Kakugawa, N. Hino, N. Hara, K. Mori, T. Tominaka, M. Kitamura, N. Maki, H. Kawano, H. Takeshima and T. Honmei
1301
1305
xxiv
Contents
Magnetizing properties of NbTi/Nb/Cu multilayer composite tubes H. Takeshima, H. Kawano, T. Hommei, N. Hara and N. Maki
1309
Reduction rate of magnetic field generated by magnetized NbTi/Nb/Cu multilayer composite tubes H. Kawano, H. Takeshima, T. Hommei, N. Hara and N. Maki
1313
Dependence of quench current and AC losses on twist directions and annealing of AC superconducting cables T. Taniguchi and R. Takahashi
1317
Recent progress of NbTi 46.5 Wt% superconductor strands at GEC alsthom C. E. Bruzek, P. Moca~'r, P. Sulten, F. Peltier, C. Kohler, E. Sirot and G. Grunblatt
1321
A new concept for the composition of oxide superconducting wires for Ac windings operating at LN2 temperature M. Iwakuma, K. Funaki, M. Takeo, K. Yamafuji, M. Konno, Y. Kasagawa, K. Okubo, L Itoh and S. Nose
1325
AC loss and current distribution in parallel conductors for Bi2223 HTS transformer windings M. Iwakuma, K. Funaki, K. Kanegae, H. Shinohara, T. Wakuda, M. Takeo, K. Yamafuji, M. Konno, Y. Kasagawa, K. Okubo, I. Boh, S. Nose, M. Ueyama, K. Hayashi and K. Sato
1329
Mutual correlation factor based high speed quench detection K. Nakamura, T. Kasuga and Y. Abe
1333
A proposal system of monitoring of multi-coil superconducting magnets system using hierarchy fuzzy theorem Y. Uriu, A. Ninomiya, Y. Kanda, T. Ishigohka, J. Yamamoto and T. Mito
1337
A typical velocity for heat destruction and restoration of S-state of short sample HTSC film M. Lutset
1341
Oxide Superconductors Tapes and Wires Bismuth-based HTS wires and their application progress K. Sato, K. Hayashi, K. Ohkura, K. Ohmatsu, T. Hikata, T. Kaneko, T. Kato, T. Sashida, M. Ueyama, J. Fujikami, M. Ito, S. Kobayashi, N. Saga, S. Hahakura, T. Shibata, T. Masuda and S. Isojima
1347
Improvement of electromagnetic properties of Ag-clad bi-based superconducting tapes through process control S. X. Dou, M. Ionescu, W. G. Wang, J. Horvat, Y. C. Guo, H. K. Liu, K. H. Mfiller and C. Andrikidis
1353
Influence of oxygen partial pressure on critical current density of Ag/Bi2223 tapes K. Osamura, S. Nonaka and Y. Katsumura
1357
The effect of stress on energization of the coils made of Ag sheathed and high strength Ag-Mg alloy sheathed Bi-2223 tape K. Tasaki, Y. Yamada, K. Yamamoto, H. Onoda, A. Tanaka, M. Urata, O. Horigami, T. Hasegawa, T. Koizumi, T. Kiyoshi and K. Inoue
1361
Electromagnetic strain of a BSCCO-2223 silver-sheathed tape reinforced by a stainless steel tape K. Ohkura, M. Ueyama, K. Sato, S. A waji and K. Watanabe
1365
Contents
xxv
Effects of hot press on critical currents in Ag-sheathed Bil.8Pbo.4Sr2Ca2.2Cu3Ox tapes J. Chikaba and K. Samoto
1369
Fabrication and properties of silver sheathed mono- and multi-filamentary Bi-2223 tapes X. Wu, Z. Duan, X. Tang and L. Zhou
1373
Influence of doping Ti, Zr or Hf into Ag-Cu alloy sheath of Bi-2223 tapes M. Ishizuka, Y. Tanaka, T. Hashimoto and H. Maeda
1377
Influence of the phase distribution on the properties of Ag-sheathed Bi-2223 tapes N. Futaki, S. Nagaya, T. Shimonosono, A. Kume, M. Nakagawa, N. Sadakata, T. Saitoh and O. Kohno
1381
Alloy sheath effects of (Bi,Pb)2Sr2Ca2Cu3Oy tapes H. Ikeda, Y. Tanaka, R. Yoshizaki, M. Ishizuka and H. Maeda
1385
Fabrication and properties of Ag-Mg-Ni alloy sheathed (Bi,Pb)2SrzCa2Cu3Ox tapes J. Sato, K. Nomura and S. Kuma
1389
Effect of Ca2PbO4 in precursor powder on the heat treatment of Bi2223/Ag superconducting tapes Y. C. Guo, W. G. Wang, H. K. Liu and S. X. Dou
1393
Advances in the fabrication of multifilamentary Bi(Pb)-2223/Ag tapes N. V. Vo, H. K. Liu, S. X. Dou and E. W. Collings
1397
Improvement of strong links in Ag/Bi-2223 superconducting tapes J. Horvat, Y. C. Guo and S. X. Dou
1401
Fabrication of multilayered composite wire of Bi-based superconductor M. Mimura, K. Kosugi, H. Ii, N. Uno, Y. Tanaka and K. Satou
1405
Fabrication of double-sheathed Bi-2212 multifilamentary wire T. Hase, K. Shibutani, S. Hayashi, M. Shimada, R. Ogawa, Y. Kawate, T. Kiyoshi and K. Inoue
1409
In-plane aligned YBCO thin film tape fabricated by pulsed laser deposition K. Hasegawa, N. Yoshida, K. Fujino, H. Mukai, K. Hayashi, K. Sato, T. Ohkuma, S. Honjyo, H. Ishii and T. Hara
1413
Modified solution-sol-gel process to synthesize ultrafine BiPbSrCaCuO powder with low carbon content C. B. Mao, L. Zhou, X. Z. Wu and X. Y. Sun A new chemical process to synthesize ultrafine BiPbSrCaCuO powder L. Zhou, C. Mao, X. Wu
1417
1421
PART THREE Bulk and Thin Film Superconducting properties and structures in Bi-2212 oxide synthesized by a diffusion process Y. Yamada, Y. Hishinuma, F. Yamashita, K. Wada and K. Tachikawa Influence of composition on microstructures and formation of the (Bi,Pb)-2223 phase in the partial-melting and sintering process X. Y. Lu, A. Nagata, M. Yasuda, K. Sugawara and S. Kamada Effect of silver addition on superconducting properties of Bi-2223 K. Kawasaki, H. Ikeda, R. Yoshizaki and K. Yoshikawa
1427
1431
1435
xxvi
Contents
The effect of the nominal composition On seeded melt grown Y-Ba-Cu-O crystals S. Takebayashi, I.-Y. Sang and M. Murakami
1439
Effect of Pt addition on Y-Ba-Cu-F-O superconductors T. Akune, K. Mizusaki, S. Iwaski, N. Sakamoto, T. Hamada and T. Ogushi
1443
Kinetics of YBa2Cu3Ox formation and crystal defects in PMP-processed superconductor L.-A. Zhou, K.-G. Wang, P.-X. Zhang, P. Ji, X.-Z. Wu, S.-K. Cheng
1447
Synthesis of 90 K superconductor Yo.9Cao.~Ba2Cu4Os at ambient pressure X. G. Zheng, H. Kuriyaki, M. Suzuki, M. Taira, C. Xu and K. Hirakawa
1451
New process to control critical currents of NdBa2Cu307_A Y. Shiohara, M. Nakamura, T. Hirayama, Y. Yamada and Y. Ikuhara
1455
Microstructure and superconductive properties of NdBa2Cu3Ox superconductor prepared by floating zone melting method S. Matsuoka, M. Sumida, T. Umeda and Y. Shiohara
1459
Critical-current densities of twin-free Nd(Ba~_x,Ndx)2Cu307_A single crystals grown by the traveling-solvent floating-zone method T. Egi, J.-G. Wen, T. Machi, K. Kuroda, H. Unoki and N. Koshizuka
1463
NdBa2Cu307_A single crystal growth by the traveling-solvent floating-zone method K. Kuroda, I.-H. Choi, T. Egi, H. Unoki and N. Koshizuka
1467
Improvement in Ic-B performance of Tl-base high-Tc oxides prepared by a diffusion process K. Tachikawa, A. Kikuchi and M. Ogasawara
1471
The effects of V2O 5 addition on the superconducting properties of YBa2Cu307.A Z.-Q. Yang, X.-Z. Su, H. Tang, D. Zeng, C. Zhang, Y.-Z. Wang and G.-W. Qiao
1475
Influence of the substrate on the anisotropy of the critical current in sputtered YBCO films O. Sarrhini, J. Baixeras and A. Kreisler
1479
Effect of early stage deposition conditions on properties of Y-Ba-Cu-O thick films prepared by laser ablation K. Shingai, Y. Yamagata, T. Ikegami and K. Ebihara
1483
Significant improvement in the surface morphology of off-center magnetron sputtered EuBa2Cu3Ov_A films on MgO K. Tsuru, S. Karimoto, S. Kubo and M. Suzuki
1487
Fabrication of Hg-Ba-(Ca)-Cu-O thin films by DC sputtering S. Koba, T. Ogushi, S. Higo, Y. Hakuraku and I. Kawano
1491
Preparationof BiSrCaCuO thin films on
1495
Improvement of magnetic shielding effect for an YBCO thick-film cylinder by using a hybrid ferromagnetic cylinder M. Itoh, F. Pavese, M. Bianco, M. Vanolo, K. Mori and T. Minemoto
1499
R-plane sapphire substrates with MgO buffer layer Y. Hakuraku, K. Maezono, H. Ueda, S. Koba, S. Higo and T. Ogushi
Magnetic Properties Flux pinning and flux dynamics in high-T~ superconductors T. Fujiyoshi, T. Sueyoshi, T. Yatsuda and R. Nishihara
1505
Entrapment of Y211 particles and Jc properties in melt-processed Y123 crystals A. Endo, H. S. Chauhan and Y. Shiohara
1509
Contents Thermomagnetic behaviors related to pulse excitation for HTS permanent magnets
xxvii 1513
M. Qiu, S. Han and L.-Z. Lin
Effect of crossed flux on magnetization and magnetization dynamics in melt-textured Y1BazCu3OT_A
1517
S. Manzoor and S. K. Hasanain
Electronic structure calculations of impurities pinning magnetic flux lines in oxide superconductors
1521
K. Fukushima
Flux-line depinning of vibrating ceramic YBa2Cu307_A
1525
Y. Fujii, H. Hamada, K. Takahashi, Y. Yamamoto, T. Shigematsu, M. Nakamura, M. Yamaguchi and T. Shigi
Critical state and magnetization property of thin YBCO films in perpendicular magnetic fields
1529
Y. Mawatari, A. Sawa, H. Obara and H. Yamasaki
Surface morphology of oxygen-deficient YBa2Cu307_Athin films
1533
G. Samadi Hosseinali, Y. Meslmani, H. W. Weber, E. Stangl, S. Proyer and D. Bduerle
Magnetic response of a single-crystal YBa2Cu3Oy thin film
1537
H. Yasuoka, S. Tochihara, H. Mazaki, M. Komatsu and M. Nagano
Intergranular coupling and secondary phases in (Bi2_x,Pbx)Sr2Ca2Cu3Olo+A/Ag tapes
1541
C. M. Friend, L. Le Lay, T. P. Beales, M. Penny and C. Beduz
Intergranular magnetization of Pb-Bi-Sr-Ca-Cu-O/Ag tapes
1545
C. Andrikidis, K.-H. Mfiller, H. K. Liu and S. X. Dou
Intergranular and intragranular AC hysteresis losses in monofilamentary Pb-Bi-Sr-Ca-Cu-O/Ag tapes
1549
K.-H. Mfiller, C. Andrikidis, H. K. Liu and S. X. Dou
Flux flow and flux creep for circulating current in Bi(2223) and Bi(2212) cylindrical bulk superconductors
1553
T. Yasunaga and D. Ito
Critical currents and irreversibility in (Bi,Pb)2Sr2Ca2Cu3Ox/Ag tape
1557
K. Zadro, E. Babi(, I. Kt~sevi(, J. Horvat, H. K. Liu and S. X. Dou
Reversible fluxoid motion in superconducting fullerene Rb3C60
1561
N. Sakamoto, Y. Ide, T. Akune, M. Baenitz and K. Liiders
Transport Properties On the scaling of the transport characteristics in high Tc superconductors from the view point of pin-fluctuation
1567
T. Kiss, T. Nakamura, M. Takeo, F. Irie and K. Yamafuji
Transport properties of Bi-2212 single crystals
1571
T. Tsukamoto, H. Andoh, T. Sugiura, G. Triscone, J.-Y. Genoud, E. Walker and N. Hase
Electrical properties of Ag sheathed Bi-2223 tapes
1575
H. Ii, A. Kimura, M. Mimura, N. Uno, Y. Tanaka, H. Ishii, S. Honjo and T. Hara
Irreversible properties of Tl-based high-Tc superconductors G. Brandstdtter, G. Samadi Hosseinali, K. Kundzins, Y. Meslmani, F. M. Sauerzopf, W. Stra!f, B. Starchl and H. W. Weber
1579
xxviii Contents Scaling behavior of current-voltage characteristics in single- and poly-crystalline YBCO films H. Ishii, S. Singo, S. Harada, S. Hirano, N. Yoshida, K. Fujino, K. Sato and T. Hara
1583
Temperature and magnetic-field scaling of transport characteristics in high quality Y IBa2Cu307_A thin films T. Nakamura, T. Kiss, M. Takeo and K. Yamafuji
1587
Superconductivity of YBa2Cu3Ox on Lao.7Cao.3MnO3 T. Yotsuya, Y. Kakehi and S. Ogawa
1591
AC Loss Self field AC transport properties and losses in high-Tc superconducting wires H. Ishii, J. Fujikami, S. Hirano, N. Yoshida, K. Sato and T. Hara
1597
AC losses in silver sheathed Bi-2223 wires M. Nakagawa, S. Nagaya, T. Shimonosono, A. Kume, N. Futaki, N. Sadakata, T. Saitoh and O. Kohno
1601
AC loss characteristics of Ag-sheathed Bi-2223 tapes in transverse external magnetic field M. Sugimoto, A. Kimura, M. Mimura, Y. Tanaka, H. Ishii, S. Honjo and T. Hara
1605
Eddy current and creep-based losses in CTFF-type Bi:2223/Ag multifilamentary tapes M. D. Sumption, N. V. Vo, S. X. Dou and E. W. Collings
1609
AC power loss of Bi-2223 tapes in self-field S. Fleshier
1613
Analysis of AC losses in high Tc superconducting wires N. Amemiya, T. Yamazumi and O. Tsukamoto
1617
A sensitive calorimetric AC-loss measurement technique for high-Tc superconducting wires J. R. Cave, P. Dolez, M. Aubin, D. W. A. Will~n and R. Nadi
1621
Other Properties Electronic properties of Hg-1201 oxide superconductors determined by NMR and STM investigations W. Hoffmann, H. Breitzke, M. Baenitz, M. Heinze, K. Lfiders, A. A. Gippius, E. V. Antipov, P. Jess, U. Hubler, H. P. Lang and H.-J. Gfintherodt
1627
Characterization of femtosecond time-transient nonequilibrium state in YBCO thin films M. Tonouchi, M. Tani, Z. Wang, K. Sakai, S. Tomozawa, M. Hangyo and Y. Murakami
1631
The research of thermal expansion of high-Tc-superconductors L. Xu, S.-L. Zhou, J. Ye and J.-M. Xu
1635
Tunneling spectroscopy in high T¢ Y-system ceramic superconductor M. Suzuki, M. Taira, X.-G. Zheng and T. Hoshino
1639
Characterization of intrinsic Josephson junction made of 60 K phase YBCO thin films M. Tonouchi, A. Itoh, T. Yasuda, H. Shimakage, Z. Wang and S. Takano
1643
Reentrant transport behavior in CaLaBaCu3Oy M. Akinaga
1647
Application of 3D-Josephson junction array model to critical temperature of high T¢ oxide superconducting under uniaxial pressure and under hydrostatic pressure C. Kawabata
1651
Contents Flexural strength and fracture toughness of YBaCuO superconductors
xxix 1655
H. Fujimoto and T. Ban
Metallic Superconductors Nb3AI Effect of static and cyclic strain on Jc of Nb3A1 CIC conductors: A comparison with Nb3Sn
1661
W. Specking, T. Ando, H. Tsuji, A. Mikumo and Y. Yamada
Experimental result of 13T-46kA Nb3A1 conductor in SULTAN
1665
Y. Nunoya, M. Oshikiri, T. Ando, Y. Takahashi, H. Nakajima, M. Sugimoto, H. Tsufi, S. Shimamoto, Y. Yamada, M. Konno, B. Blau, I. Rohleder, G. Vecsey, S. Shen and H. Katheder
Critical current density of Nb3A1 superconducting wire for high field magnet by rapid-quenching process
1669
K. Fukuda, G. Iwaki, M. Kimura, S. Sakai, Y. Ioima, T. Takeuchi, K. Inoue, N. Kobayashi, K. Watanabe and S. A waji
Development of practical Nb3A1/Cu multifilamentary superconducting strand
1673
A. Mikumo, N. Ayai, Y. Yamada, K. Takahashi, K. Sato, N. Koizumi, M. Sugimoto, T. Ando, H. Tsuji and S. Shimamoto
Improvement of critical current density and residual resistivity on jelly-roll processed Nb3AI superconducting wires
1677
N. Ayai, A. Mikumo, Y. Yamada, K. Takahashi, K. Sato, N. Koizumi, T. Ando, M. Sugimoto, H. Tsuji and S. Shimamoto
Fabrication and superconducting properties of Nb3A1 composite wire using the jelly roll process
1681
Y. Sakagami, T. Yamazaki, N. A oki, M. Ichihara, T. Masegi, S. Murase, K. Matsui, K. Ushigusa, M. Kikuchi and Y. Takahashi
Development of multifilamentary Nb3A1 superconducting wires by the jelly-roll method
1685
K. Aihara, Y. Hanaoka, T. Suzuki, T. Nabatame, G. Iwaki, K. Fukuda, S. Sakai, K. Sasaki and K. Kikuchi
Superconducting properties and microstructure of Nb3A1 wire fabricated by the jelly-roll method
1689
Y. W. Hanaoka, T. Suzuki, T. Nabatame, K. Aihara, G. Iwaki, S. Sakai and K. Kikuchi
Superconducting properties of Nb3A1 by reaction between a-phase Nb2A1 and Nb powders
1693
N. Harada, Y. Fukuda, H. Ichikohara, K. Osaki, N. Tada, S. Sakai and K. Watanabe
Critical current density characteristics of Nb3A1 multifilamentary wires continuously fabricated by rapid-quenching
1697
Y. Iijima, M. Kosuge, T. Takeuchi and K. Inoue
Nb3A1 thin films made by RF magnetron sputtering process with a single target
1701
K. Agatsuma, H. Tateishi, K. Arai, T. Saitoh, N. Sadakata and M. Nakagawa
Nb3Sn Development of the alumina-copper reinforced Nb3Sn wire for coil fabrication S. Murase, S. Nakayama, K. Koyanagi, T. Masegi, S. Nomura, M. Urata, K. Shimamura, K. Amano, N. Shiga, K. Watanabe and N. Kobayashi
1707
xxx
Contents
Effects of the strain on the critical current of the powder metallurgy processed Nb3Sn superconducting wires N. Matsukura, T. Miyazaki, Y. Inoue, T. Miyatake, M. Shimada, R. Ogawa and M. Chiba
1711
Improvements in the properties of internal-tin Nb3Sn strands E. Gregory, E. Gulko, T. Pyon and L. F. Goodrich
1715
Critical current density at high fields for bronze-processed (Nb.Ti.Ta)3Sn superconducting wires G. Iwaki, M. Kimura, S. Sakai, N. Kobayashi, K. Watanabe and S. A waji
1719
Mechanical and superconducting properties of multifilamentary Nb3Sn wires with CuNb reinforcing stabilizer S. Iwasaki, K. Goto, N. Sadakata, T. Saito, O. Kohno, S. A waji and K. Watanabe
1723
The effects of coated film on (NbTi)3Sn wire on friction and electric contact resistance A. Iwabuchi, H. Funayama, T. Shimizu, M. Ono and T. Hamashima
1727
Manufacture of Nb3Sn strands for ITER by internal tin diffusion process K. Egawa, Y. Kubo, T. Nagai, M. Wakata, F. Uchikawa, 0. Taguchi, K. Wakamoto, M. Morita, T. Isono, Y. Nunoya, K. Yoshida, M. Nishi and H. Tsuji
1731
Development of NbTi and Nb3Sn conductors for 1 GHz NMR spectrometer K. Itoh, M. Yuyama, T. Kiyoshi, T. Takeuchi, K. Inoue, H. Maeda, T. Miyatake and M. Shimada
1735
Comparison of properties between bronze-processed Nb3Sn wires with different tin content T. Miyazaki, T. Miyatake, M. Shimada, H. Kurahashi, I. Tatara and N. Kobayashi
1739
Reduction of critical current in Nb3Sn cables as a function of transverse p r e s s u r e - A comparison with a finite element model J. M. van Oort, R. M. Scanlan and H. H. J. ten Kate
1743
Other Materials Superconductivity of reactively sputtered TaN film for ULSI process H. Kubota, K. Wakasugi, M. Tokunaga, T. Sumita and M. Nagata
1749
Preparation of 1 m long superconducting NbN tapes by DC magnetron sputtering M. Suzuki, T. Kiboshi and K. Sugawara
1753
SN-interface flux pinning in bronze processed V3Siconductor Y. Nemoto, T. Takeuchi, H. Maeda and K. Togano
1757
A C Loss and Electromagnetic Properties AC losses of proximity-induced superconducting Cu in NbTi filamentary composite T. Moriya, H. Kubota, K. Yasohama, Y. Kubota and T. Ogasawara Suppression of eddy current loss in bare-copper Rutherford cables using stainless steel cores of various thicknesses E. W. Collings, M. D. Sumption, S. W. Kim, M. Wake, T. Shintomi and R. M. Scanlan
1763
1767
Improvement of hysteresis-losses in internal tin diffusion processed Nb3Sn wires Y. Kubo, K. Egawa, T. Nagai, M. Wakata, F. Uchikawa, O. Taguchi, K. Wakamoto, M. Morita, T. Isono, Y. Nunoya, K. Yoshida, M. Nishi and H. Tsuji
1771
AC losses of Nb3Sn AC multifilamentary superconducting wires due to transport current S. Fukui, M. Ito, N. Amemiya, O. Tsukamoto and M. Hakamata
1775
Contents
xxxi
Proximity-effect-induced filament coupling in NbTi wires K. Yasohama, N. Azuma, Y. Kubota and T. Ogasawara
1779
Exploration of micro-bridging effect in Nb3Sn superconductors D. Mao, M. Yuyama, K. Itoh, H. Wada and Y. Murakami
1783
Round robin test for the method for critical current measurement of Nb3Sn composite superconductors K. Itoh, Y. Tanaka and K. Osamura
1787
Standardization of the method for the room temperature tensile test of Cu/Nb-Ti composite superconductors S. Sakai, K. Osamura, M. Hojo, T. Ogata and M. Shimada
1791
Standardization of the method for the determination of the residual resistance ratio (RRR) of Cu/Nb-Ti composite superconductors S. Murase, T. Saitoh, T. Matsushita and K. Osamura
1795
Standardization of the test method for critical current measurement of Cu/Cu-Ni/Nb-Ti composite superconductors R. Ogawa, Y. Kubo, Y. Tanaka, K. Itoh, K. Ohmatsu, T. Kumano, S. Sakai and K. Osamura
1799
Third harmonic AC susceptibility in multifilamentary wires E. S. Otabe, R. Kitamura and T. Matsushita
1803
RRR evaluation of niobium using AC susceptibility measurement M. Wake and K. Saito
1807
Pinning and Stability Orientational dependence of critical current density of NbTi thin film in a magnetic field M. Takeda, K. Nishigaki and H. Toda Computer simulations of vortex pinning properties by line pins O. Ichimaru and E. Kusayanagi Flux pinning properties of ultrafine multifilamentary NbTi superconducting wires with Nb island-type artificial pins O. Miura, C. Tei, D. Ito and S. Endo
1813
1817
1821
Evaluation of recovery current of the helical coil for LHD S. Imagawa, N. Yanagi, T. Mito, T. Satow, J. Yamamoto, O. Motojima and the LHD Group
1825
Delayed ramps in composites described in a two-layer model A. Akhmetov, M. Takeo and T. Imayoshi
1829
Normal zone at the edge of a flat two-layer superconducting cable A. A. Akhmetov, K. Kuroda, H. Tanaka and M. Takeo
1833
Statistical analysis of scattering in quench current of ac superconducting wires N. Hayakawa, M. Wakita, M. Hikita and H. Okubo
1837
Minimum quench energy and normal zone propagation velocity of bronze processed (Nb,Ti)3Sn superconducting wire with CuNb reinforcing stabilizer T. Onodera, K. Watanabe, K. Noto, T. Sugiura, M. Matsukawa, S. Awaji, K. Goto, S. Iwasaki, N. Sadakata, O. Kohno and J. Yamamoto
1841
xxxii
Contents
Applications R&D on NbTi and Nb3Sn superconductors for AC use in the super-GM M. Chiba, M. Hakamata, H. Chiba, N. Yoshida and T. Ichikawa
1847
Nb-Ti superconducting strands with Cu-Ni-Mn matrix and 2kA-class cables for AC use K. Miyashita, K. Sugiyama, S. Sakai, K. Kamata and H. Chiba
1851
Stability and AC losses of l kA class multifilamentary Nb3Sn superconducting cables for AC applications K. Ohmatsu, H. Yumura, K. Takahashi, K. Sato and N. Yoshida Development of kA class multifilamentary Nb-Ti superconducting cables for AC applications K. Ohmatsu, H. Yumura, K. Takahashi, K. Sato and N. Yoshida Some properties of bronze process Nb3Sn superconducting wires with Cu-Sn-Ge matrix and Nb-Ta cores for AC use S. Sakai, K. Miyashita, K. Sugiyama, K. Kamata, K. Tachikawa and H. Chiba Development of in-situ Nb3Sn wires for AC use H. Fuji, K. Goto, N. Sadakata, T. Saito, O. Kohno and J. Yoshitomi Development of industrially produced composite quench heaters for the LHC superconducting lattice magnets B. Szeless, F. Calvone and F. Rodriguez Mateos
1855
1859
1863
1867
1871
Metallic Materials Mechanical Properties Fracture toughness of SUS 316 and weld joint at cryogenic temperature A. Nishimura, R. L. Tobler and J. Yamamoto Effects of pre-cracking condition on fracture toughness of SUS 304L steel in an 8 tesla magnetic field at 4K K. Shibata, T. Shimonosono and T. Tanaka
1877
1883
Cryogenic effects on the fracture behavior of forged JJ1 type austenitic stainless steel plate Y. Shindo, K. Horiguchi, K. Sanada, T. Kobori, H. Nakajima and H. Tsuji
1887
The 4 K tensile and fracture toughness properties of a modified 316LN conduit alloy R. P. Walsh, L. T. Summers and J. R. Miller
1891
4 K mechanical properties of pure titanium for the jacket of Nb3Sn superconductors H. Nakajima, Y. Nunoya, K. Takano, T. Ando, H. Tsuji, S. Konosu, O. Ivano, T. Nakaniwa, T. Horiya and S. Ohkita
1895
Load-control tensile test of steels and alloys in liquid helium T. Ogata and T. Yuri
1899
Residual stresses in superconducting jackets after compaction F. M. G. Wong, N. A. Mitchell, M. M. Morra, P. H. Titus and R. N. Randall
1903
Vickers hardness properties of metals at cryogenic temperatures Y. Yoshino, T. Abe, A. Iwabuchi, A. Chiba and T. Shimizu
1907
An aluminium alloy for the coil clamping collars of the prototype LHC dipole magnets T. Kurtyka, M. Bona, S. Sgobba, S. Marque and B. Skoczen
1911
Contents Evaluation tests of structural materials used for liquid hydrogen storage and transportation system (Study on low temperature materials used in WE-NET, 1) T. Horiya, N. Yamamoto, T. Iida, A. Yamamoto, S. Okaguchi, N. Yaegashi, T. Doko, M. Saito, K. Yokogawa and T. Ogata Development of materials testing equipment in high-pressure hydrogen at cryogenic temperatures and effect of temperature on hydrogen environment embrittlement of steels G. Han, J.-H. He, S. Fukuyama and K. Yokogawa Hydrogen embrittlement of austenitic stainless steels at low temperature (Study on low temperature materials used in WE-NET 2) S. Okaguchi, T. Horiya, K. Ishige, M. Saito, N. Yaegashi, A. Yamamoto, H. Nakagawa, K. Yokokawa and T. Ogata
xxxiii 1915
1919
1923
Physical Properties Absolute measurements of linear thermal expansion coefficients of copper SRM 736 and some commercially available coppers in the temperature range 20 - 300 K M. Okaji, N. Yamada, H. Kato and K. Nara
1929
Possibility of AC applications of composite high-purity aluminum wires L. I. Chubraeva, S. N. Pylinina, V. E. Sigaev and V. A. Tutaev
1933
Development and investigation of composite high-purity aluminum wires L. I. Chuhraeva, V. E. Sigaev, E. A. Phtshkin, N. 1. Sahmin and E. N. Aksenova
1937
Temperature structures at high Hall drift in aluminum V. R. Sobol, O. N. Mazurenko and A. A. Drozd
1941
Inhomogeneity and local negative magnetoresistance of aluminum V. R. Sobol, O. N. Mazurenko and A. A. Drozd
1945
Magnetic field of cyclic current in cylinder conductors B. B. Boiko, V. R. Sobol and O. N. Mazurenko
1949
Magnetism due to Hall current in aluminum and copper conductors B. B. Boiko, V. R. Sobol, O. N. Mazurenko and A. A. Drozd
1953
Effects of high magnetic field and tensile stress on martensitic transformation behavior and microstructure at 4.2 K in Fe-Ni-C shape memory alloys H. Ohtsuka, G. Ghosh, K. Nagai, H. Kitaguchi and M. Uehara
1957
Analysis of temperature rises in point and line contact slidings in liquid helium I. Nitta, A. Iwabuchi, T. Takao and M. Minami
1961
Round robin tests on frictional characteristics of cryogenic materials T. Takao, A. Iwabuchi, M. Minami and I. Nitta
1965
Galvanic chrome coating of copper wire for the ITER program Y. Ipatov, V. Sytnikov, A. Rychagov and G. Svalov
1969
Non Metallic Materials Mechanical Properties Overview of the fatigue behaviour of fiber composites at low temperatures G. Hartwig and R. Hiibner
1977
xxxiv
Contents
Nonlinear fracture behavior of G-10 woven glass-epoxy laminates at liquid nitrogen temperature K. Sanada, Y. Shindo and K. Horiguchi
1981
Effects of friction on shear/compressive strength of high-density GFRP at low temperature T. Suzuki, S. Usami, H. Tsukamoto and M. Otsuka
1985
Effect of contiguity of glass bundles on the inter-laminar shear strength changed of ILSS induced by low-temperature electron irradiation Y. Tsukazaki, S. Nishijima, T. Nishiura, T. Okada and K. Asano
1989
Frictional characteristics of high strength polyethylene fiber reinforced plastics T. Takao, T. Nemoto, H. Konda, T. Kashima and A. Yamanaka
1993
Cryogenic material properties in tension and shear of reactor irradiated GFRPs K. Humer, S. M. Spiebberger, H. W. Weber, E. K. Tschegg, N. A. Munshi and P. E. Fabian
1997
Shear compression strength of organic insulation systems after reactor irradiation at 4 K H. Gerstenberg, E. Kriihling, H. Katheder, R. Maix and M. SO'll
2001
Scaling experiments and mechanical properties of CFRPs at 293 K and 77 K after room temperature reactor irradiation S. M. Spieflberger, K. Humer, E. K. Tschegg, H. W. Weber, S. Valthe, K. Noma and IT. Iwasaki
2005
Properties of candidate insulation systems for ITER CS model coils at low temperatures H. Mitsui, H. Hirai, H. Moriyama, Y. Sumiyoshi, T. Sasaki, M. Sugimoto, T. Ando, H. Tsuji and R. Vieira
2009
Enhanced creep of epoxy-resin during electron beam irradiation T. Nishiura, S. Nishijima, T. Okada and Y. Tsukasaki
2013
Possible hazards following irradiation of superconducting magnet insulation D. Evans, R. P. Reed and N. J. Simon
2017
Physical Properties Low-temperature thermal properties of amorphous Polycarbonat and Polystyrene M. Jiickel, F. von Schoenebeck, U. Escher and A. Gladun
2023
Experimental measurement on the specific heat of highly porous materials W. Shi, J.-Y. Wu, R.-Z. Wang and Z.-M. Tang
2027
Low-temperature thermal conductivity of amorphous Polycarbonat under high pressure M. Jd'ckel, F. Weise and R. Geilenkeuser
2031
Composite matrix for low thermal contraction down to cryogenic temperature K. Nojima, T. Ueki, K. Asano, S. Nishijima and T. Okada
2035
Functional Materials Mechanical and thermal properties of zirconia at cryogenic temperature S. Ueno, S. Nishijima, A. Nakahira, T. Sekino, T. Okada and K. Niihara
2041
Strength characteristics of alumina ceramics at cryogenic temperatures N. Suzuki, K. Suzuki, T. Murakami and T. Uchida
2045
Residual strain of pipes composed with high strength polyethylene fiber reinforced plastics at cryogenic temperature T. Kashima, A. Yamanaka, S. Nishijima and T. Okada
2049
Contents
xxxv
Advanced composite materials for cryogenic support structures R. K. Giesey
2053
Fracture toughness of epoxy filled with nano-scale silica at cryogenic temperature S. Ueno, S. Nishijima, M. Hussain, A. Nakahira, F. Sawa, T. Okada and K. Niihara
2057
Improvement of fracture toughness of epoxy resin for cryogenic use T. Ueki, K. Nojima, K. Asano, S. Nishijima and T. Okada
2061
Magnetothermal properties of amorphous Gd7oCu3oand Dy7oCu3oalloys X.-Y. Liu, J. A. Barclay, R. B. Gopal, M. Ffldegtki, R. Chahine and T. K. Bose
2065
Crystal axis dependence of magneto-thermal properties in RA103 (R: Dy, Ho and Er) single crystals using for magnetic refrigeration H. Kimura, T. Numazawa, M. Sato, T. lkeya, T. Fukuda and K. Fujioka
2069
Magnetic field dependence of thermal conductivity in rare-earth oxides for heat switch application T. Numazawa, H. Kimura, M. Sato, A. Sato, K. Shimamura and T. Fukuda
2073
Author Index
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Organized under the auspices of International Cryogenic Engineering Committee - ICEC International Cryogenic Materials Committee- ICMC International Institute of Refrigeration, Commission A 1/2 - IIR We acknowledge the support and sponsorship of Cryogenic Association of Japan Kitakyushu City Association for Promotion of Electrical Electronic and Information Engineering Commemorative Association for the Japan World Exposition (19 70) Kyushu Bureau of International Trade and Industry Fukuoka Prefecture Fukuoka Science & Technology Foundation ORGANIZING COMMITTEES INTERNATIONAL CRYOGENIC ENGINEERING COMMITTEE P. Komarek (Chairman) (Vice Chairman) T. Mitsui C. Rizzuto (Vice Chairman) P. Berglund A.F. Clark G. Gistau H. Hirabayashi G. Klipping (Honorary Member) P. Lebrun L.Z. Lin K.G. Narayankhedkar H. Quack H. Rogalla R.S. Safrata R.G. Scurlock A.J. Steel VoV. Sytchev S. Van Sciver M. Wood (Honorary Member) K. Yamafuji K.A. Yushchenko L. Zhang INTERNATIONAL CRYOGENIC MATERIALS COMMITTEE E.W. Collings (Chairman) A.F. Clark D. Evans F.R. Fickett H.C. Freyhardt G. Hartwig K.T. Hartwig Y.Y. Li H. Maeda T. Okada R.P. Reed D.B. Smathers L.T. Summers K. Tachikawa H.H.J. Ten Kate G.R. Wagner H.W. Weber K.A. Yushchenko L. Zhou xxxvii
LOCAL ORGANIZING COMMITTEE
T. Mitsui (Chairman) K. Yamafuji (Vice Chairman) M. Akiyama Y. Akiyama A. Hikita H. Hirabayashi J. Hosoda Y. Kawate K. Kitazawa S. Kosaka N. Koshizuka H. Maeda T. Matsushita H. Nakashima K. Nagata Y. Nakano N. Noda O. Ogino H. Ogiwara T. Okada S. Shimamoto T. Shintomi K. Tachikawa M. Takeo T. Tashiro O. Tsukamoto J. Yamamoto K. Funaki (Conference Secretariat)
LOCAL PROGRAMME COMMITTEE
(ICEC) H. Hirabayashi(Chairman) O. Tsukamoto (Vice Chairman) N. Amemiya T. Haruyama D. Itoh H. Maeda Y. Matsubara M. Nishi Y. Okuda Y. Takata
(~CMC)
T. Okada (Chairman) H. Maeda (Vice Chairman) K. Enpuku T. Hashimoto S. Kosaka T. Matsushita S. Nishijima T. Ogata M. Suzuki
INDUSTRIAL EXHIBITION COMMITTEE
M. Takeo (Chairman) T. Itoh Y. Kamioka T. Noguchi
xxxviii
EXHIBITORS Aisin Seiki Co. Ltd Arisawa Mfg Co. Ltd Cryogenic Association of Japan, Refrigeration Commission Dowa Mining Co. Ltd Engineering Research Association for Superconducting Generation Equipment and Materials (Super-GM) Fuji Electric Co. Ltd Fujikin Incorporated The Furukawa Electric Co. Ltd Gec Alsthom Hitachi Cable Ltd Hitachi Ltd International Superconductivity Technology Center, Superconductivity Research Laboratory Ishikawajima-Harima Heavy Industries Co. Ltd Iwatani International Corporation /Iwatani Plantech Corporation Japan Atomic Energy Research Institute Japan High Tech Co. Ltd Kabelmetal electro GmbH Kobe Steel Ltd Koike Sanso Kogyo Co. Ltd Kyushu Electric Power Co. Inc Kyushu University, Research Institute of Superconductivity Linde AG. Lydall Corporation Mitsubishi Electric Corp Nagase & Co. Ltd National Institute for Fusion Science National Laboratory for High Energy Physics, KEK Niki Glass Co. Ltd Nippon Automatic Control Company Nippon Sanso Corporation Oxford Instruments Ltd Railway Technical Research Institute Sumitomo Heavy Industries Ltd Suzuki Shokan Co. Ltd Taiyo Toyo Sanso Co. Ltd Teisan K.K. Tokyo Electric Power Company Toshiba Corporation Weka AG.
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Foreword
The joint conference of the 16th International Cryogenic Engineering Conference (ICEC) and 1996 International Cryogenic Materials Conference (ICMC) was held in Kitakyushu, Japan during 20-24 May 1996. For this joint conference, a Local Organizing Committee, a Local Programme Committee and an Industrial Exhibition Committee took care well for preparation and management of the conference under the strong support of the Cryogenic Association of Japan (CAJ). The conference was also co-sponsored by the International Institute of Refrigeration (IIR) -Commission A1/2. The conference was the largest so far with over 700 registered delegates from 21 countries. The scientific programme consisted of nearly 490 papers for oral (110) and poster (380) presentation accompanied by an interesting and innovating Industrial Exhibition of cryogenic equipment by 39 exhibitors. Seven plenary invited talks gave overviews to light up the most significant topics to the progress of Cryogenic Engineering and Cryogenic Materials. All the papers in this proceedings have been reviewed by the authorized review coordinators nominated by the Local Programme Committee. This proceedings is not a supplement issue of Cryogenics, however, in order to keep good relationship between the conference and the journal, we asked the session chairmen to recommend high quality papers for Cryogenics. We hope that such papers with amplified and extended contents will soon appear in Cryogenics. During the conference, the International Cryogenic Engineering Committee presented the 1996 Mendelssohn Award to Prof. O.V. Lounasmaa of the Helsinki University for his great contributions to cryogenics society. It is our pleasure to inform all of you that the next conference, ICEC17, will be held from 14 to 17 July 1998 in Bournemouth, England under the chairmanship of Professor R.G. Scurlock, Institute of Cryogenics, Southampton University.
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.,
T. Mitsui
(Chairman of the Local Organizing Committee)
K. Yamafuji (Vice-chairman of the Local Organizing Committee)
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INTERNATIONAL
CRYOGENIC
ENGINEERING
COMMITTEE Chairman: Prof. Dr. P. Komarek KfK-ITP Postfach 3640 ,., . . . . . ,ors, . . . . . Germany Tel.: 49-7247-823500 Fax.: 49-7247-822849
Secretary: Dr. H. Quack Bodenackerstrasse 48 CH-8330 Pf&ffikon ZH Switzerland Tel.: 41-1-950 3110 Fax.: 41-1-951 0221
D-76021 Karlsruhe
Dear Conference Participants It is a groat honour and pleasure for me to welcome you on behalf of the International Cryogerlic Engineering Committee to this conference. After having the conference two times in Europe and before in China, it came back to Japan now, one of its birth places. Here this conference serie has already been hosted with an extraordinary spirit of hospitality for three times. Being again in Japan is well justified, if one regards the amount of contributions by Japanese scientists and engineers to the progress in superconductivity and cryogenics. It also became a good tradition to hold ICEC from time to time jointly with ICMC, recognizing the need for a joint effort of material specialists and project specialists to proceed with applications of superconductivity and cryogenics. I hope, that at this time again the discussions among both societies will be vital and fruitful for a mutual benefit. The foundation for the success of this conference is layed by your participation and submitted contributions, more than 600, forming the seed for a rich scientific harvest. In that spirit, let me thank you all for coming and let me encourage you to support the progress of this conference very actively. Before concluding with best wishes to you for a nice and fruitful stay in Kitakyushu, let me thank our hosts, conducted by Dr. Mitsui, Prof. Hirabayashi, Prof. Yamafuji and Dr. Okada for their excellent preparatory work to make this conference possible. I hope that those who were not able to come to Japan this time, will at least benefit from studying the Proceedings. Let me conclude in that sense with best wishes to all of you.
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._.i .
..
P. Komarek Chairman of the ICEC
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GREETINGS FROM THE CHAIRMAN OF ICMC
The International Cryogenic Materials Conference (ICMC) board together with the Cryogenic Engineering Conference (CEC) jointly holds the well-known biennial series of "CEC/ICMC" conferences, in the odd-numbered years. Between these large UnitedStates-based meetings, ICMC generally organizes topical symposia held in various parts of the world in response to the wishes of ICMC's numerous international board members. Over the years such meetings have taken place worldwide. But in place of a topical meeting ICMC has from time to time joined with the International Cryogenic Engineering Conference in the sponsoring of the joint ICEC/ICMC meetings in countries other than the US. The first of such meetings was held in Kobe, Japan, in May, 1982. The second took place ten years later in Kiev, Ukraine, in June, 1992, and we are very pleased to be holing the 1996 joint ICEC/ICMC meeting again in Japan, in the City of Kitakyushu. A splendid program of papers in cryogenic engineering, cryogenic structural materials (both metallic and nonmetallic), and low-temperature and high-temperature applied superconductivity in its various forms is offered along with a very impressive commercial exhibition of cryogenic equipment, materials, and machines. In introducing the ICEC16/ICMC conference, we must not forget that although ICEC and ICMC offer their names to it, the conference owes its success to the small army of engineers, scientists, and their assistants who make up the Local Organizing Committee (chaired by T. Mitsui and K. Yamafuji, both members of the ICEC Board), the Local ICEC Program Committee (chaired by H. Hirabayashi, ICEC Board, and O. Tsukamoto), the Local ICMC Program Committee (chaired by T. Okada and H. Maeda, both members of the ICMC Board), the Industrial Exhibition Committee (chaired by M. Takeo) other essential groups led by the Conference Secretary, K. Funaki. To all of these, the ICMC Board offers its thanks and congratulations, and trusts that the planned three-volume Conference Proceedings will provide a valuable record of the states of cryogenic engineering and materials as of May, 1996. E. W. Collings Chairman, ICMC Board
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INTERNATIONAL
CRYOGENIC
ENGINEERING
COMMITTEE Chairman Prof. Dr. P. Komarek KfK-ITP Postfach 3640 ,~ t - - D - 7 5 0 0 N a ,.... o, L,, .L_ Germany Tel. 49-7247-823500 F a x . 49-7247-822849
Secretary: Dr. H. Quack Bodenackerstrasse 48 CH-8330 Pf&ffikon ZH Switzerland Tel. 41-1-950 3110 Fax." 41-1-951 0221
D-76021 Karlsruhe
Mendelssohn Award
This award was astablished by the International Cryogonic Engineering Committee (ICEC) in memory of Kurz Mendelssohn (1906-1982), the founder of ICEC. Persons to be honoured by this award are selected based on their outstanding work in the field of cryogenic engineering, e.g. concerning new solutions to difficult problems, promotion of work in new fields and stimulating the community's interest in such fields and long-standing superior contributions to cryogenics. Earlier award winners were K. Oshima, P. Roubeau, J. Gardner, P. Mason, D. Petrac and H. Desportes. Based on the decision in the ICEC the Mendelssohn Award is presented at this conference to Prof. Olli V. Lounasmaa in recognition of this outstanding contributions over many decades to the worldwide development of cryogenics. As an excellent scientist in fields reaching from nuclear cooperative phenomena at pico-Kelvin temperatures to the use of SQUID's for studies in neuromagnetism, as a gifted teacher of young researchers at universities around the world and as organizer of several conferences he became famous in shaping the international cryogenic community and in stablishing cryogenics as an accepted field of physics and engineering. The International Cryogenic Engineering Committee is proud to present the Mendelssohn Award to Prof. O. V. Lounasmaa at the occasion of ICEC 16 in Kitakyushu.
E/
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P. Komarek Chairman of the ICEC
xlvii
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Olli V. Lounasmaa CURRICULUM VITAE BREVE May, 1996 Prof. Olli Viktor L o u n a s m a a, born August 20, 1930, earned his Bachelor's Degree in physics from the University of Helsinki in 1953 and his D. Phil. Degree from the University of Oxford in 1958. His thesis was on the thermodynamic properties of fluid 4He. After two years of research on liquid 4He at the University of Turku in Finland he became, in 1960, Resident Research Associate at the Argonne National Laboratory near Chicago. There he built a 3He refrigerator and completed an extensive series of measurements on the nuclear heat capacity of rare earth metals. During this period he also did some research on liquid 4He near the )~-transition. In 1965 Lounasmaa returned to Finland. He was appointed Professor of Technical Physics at the Helsinki University of Technology where he founded the Low Temperature Laboratory and became its director. Since then he has specialized in research at milli-, micro-, nano-, and even picokelvin temperatures. In 1970 he was appointed Professor of the Academy of Finland. He retired in January 1996. Three refrigerators, employing adiabatic nuclear demagnetization techniques for cooling, are currently operating in the Low Temperature Laboratory. In two of them, rotating 3He can be investigated between 0.4 and 3 mK: extensive NMR, superfluid flow, ion mobility, ultrasonic, and optical measurements have been carried out on vortices in rotating 3He-A and 3He-B. The third cryostat, intended for research on nuclear cooperative phenomena in metals, has two nuclear cooling stages working in series. This apparatus has held many world's low temperature records, currently 280 pK in the nuclear spin system of rhodium. The same cryostat has produced "record-high" negative spin temperatures, first in silver and then in rhodium (-750 pK). Three antiferromagnetic phases have been discovered in copper and antiferromagnetic and ferromagnetic phases in silver. Another double nuclear refrigerator, constructed in Helsinki for neutron diffraction studies at nanokelvin temperatures, was operating first at the Rise National Laboratory in Denmark and is now working at the Hahn-Meitner-Institute in Berlin; this is a joint Danish-Finnish-German research effort, partially financed by the European Community's SCIENCE program. Nuclear ordering in copper at T > 0 was studied in Rise (the three antiferromagnetic phases were confirmed) and will be investigated in silver at T > 0 and at T < 0 in Berlin. The Low Temperature Laboratory is also participating in the Spin-Muon Collaboration at CERN by being responsible for the construction of a very powerful dilution refrigerator. Much work has also been done by Lounasmaa and his students on the M6ssbauer effect, on thermometry below 1 K, and on the use of SQUIDs for low temperature and brain research. Large scale uses of superconductivity have been investigated as well, including the construction of a 0.17-T superconducting solenoid for whole-body MR imaging. A 1.5 T magnet, 2 m long and with an 0.8 m bore, was completed in 1988. A magnetically shielded room was built in 1980 into the Low Temperature Laboratory. With this facility, Lounasmaa's attention was partly turned to measurements of the weak magnetic signals produced by the human brain; in fact, neuromagnetism has now become his secondary field of interest. After several smaller instruments, a SQUID magnetometer with 24 simultaneously operating channels was put to use in 1989. A 122-SQUID device, which covers the whole head and is the first instrument in this category, has been operational in the Low Temperature Laboratory since the summer of 1992. Many extensive series of magnetoencephalographic (MEG) experiments, such as studies of basic neurophysiology, localizing epileptic foci, investigating cognitive processes, and studies of signal processing in the human brain have been carried out. In 1977 Lounasmaa worked for a year at Sacley near Paris. He has also spent 1 to 2 months as Visiting Professor in the United States, Japan, India, Germany, and Denmark. The Academic Year 1 9 8 2 - 8 3 he pursued research first at the University of California in Berkeley and later at New York University. Most of 1996 he will be working at the Hahn-Meitner-Institut in Berlin. xlix
Prof. Lounasmaa is the author of over 200 scientific papers, including a book entitled "Experimental Principles and Methods below 1 K". He has given well over 200 seminar or colloquium talks at low temperature laboratories throughout the world and attended about 100 international conferences, presenting over 40 plenary or invited papers. He has also given 16 series of lectures abroad, many of them at international schools of physics. Lounasmaa has supervised the Ph.D. theses of 40 students and written expert's opinions on the qualifications of candidates for many professorships in Finland and abroad. Lounasmaa is the Director of two Large-Scale Facilities, BIRCH in neuroand cardiomagnetism and ULTI in ultralow temperature physics, established by the European Union's Human Capital and Mobility Program in 1994. He has also been active in science policy in Finland having served, for example, in 1980 and in 1984 as Chairman of Ministry of Education working groups on basic research. At the Ministry's request he made in 1995 an evaluation of the departments of mathematics, physics, chemistry, and computer sciences in Finnish universities. In 1965 Lounasmaa was elected to the Finnish Academy of Technical Sciences, in 1969 to the Finnish Academy of Sciences and Letters (President 1992), in 1974 to the Royal Swedish Academy of Sciences, in 1976 to the Societas Scientiarum Fennica, and in 1989 to the Academia Europaea; in 1986 he became Fellow of the American Physical Society. He is also past President of the Finnish Physical Society and Honorary Member since 1990. In 1969 Lounasmaa received the Th. Hom6n Prize of Societas Scientiarum Fennica, in 1973 the Emil Aaltonen Foundation Prize, in 1978 the Finnish Cultural Foundation Prize, in 1980 Tekniska F6reningen i Finland 100-year Jubileum Medal, and in 1984 the Fritz London Memorial Award, presented at the 17th International Conference on Low Temperature Physics in Karlsruhe. In 1987 he was, together with three of his colleagues, given the Award for the Advancement of European Science by the K6rber-Stiftung of Hamburg. In 1990 he received the Finnish Government's Inventors Prize and in 1991 the "Professor of the Year" title. In 1990 he was promoted to the Degree of Doctor of Philosophy h o n o r i s c a u s a by the University of Helsinki in connection with the TT ~nlversity t s 350-year celcbrations. In 1992, he received the Degree of Doctor of Technology h o n o r i s c a u s a by the Tampere University of Technology. He is Commander 1st Class of the Order of the White Rose of Finland and Commander of the Order of the Lion of Finland. In 1993 he was honored by the Forschungspreis of the Alexander von Humboldt-Stiftung of Germany and in 1994 he was awarded the Wihuri International Prize. In 1995 he received the Finnish Government's Decoration for XXX Years of Service, the first Kapitza Gold Medal, founded by the Russian Academy of Sciences on the occasion of Pjotr Kapitza's 100th birthday, and the Italgas Prize in physics, donated by the Italgas Company and awarded in collaboration with the Academy of Sciences of Torino. He also received the Espoo medal. In 1996 he was honored by the Mendelssohn Prize awarded by the International Cryogenic Engineering Committee. "
9
Prof. Lounasmaa has participated in international collaboration among low temperature physicists. He has been President of the Commission on Cryophysics of the International Institute of Refrigeration and Chairman and Member of the Very Low Temperature Physics Commission of the International Union of Pure and Applied Physics; he has also been a member of the International Cryogenic Engineering Committee. For four years he served on the Executive Committee of the European Physical Society. In 1992 he became member of Comit6 International des Poids et Mesures. He was in 1975 President of the 14th International Conference on Low Temperature Physics (LT14) and in 1984 President of the 10th International Cryogenic Engineering Conference (ICEC-10). He was President of the 7th General Conference of the European Physical Society (EPS-7), held in Helsinki in 1987. He has organized six smaller international conferences and several winter schools with foreign participation. He has been a member of the International Committees of over 20 scientific conferences held abroad. He also belongs to the Editorial Boards of two scientific journals. Lounasmaa is married and has two daughters and four granddaughters.
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Plenary lectures
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A Way to Commercialization of High-Tc Superconductors
Shoji Tanaka Superconductivity Reseach Laboratory, International Superconductivity Technology Center 10-13 Shinonome 1-chome, Koto-ku, Tokyo 135, Japan
Since the discovery of high-Tc superconductivity in 1986, the research and developments have been made big progress. In the past ten years, we learned that the oxide superconductors have completely different characters from metallic superconductors in crystal structures, conduction properties both in normal and superconducting states and so on. Therefore, it is expected that any future applications of these materials are strongly bound by such special characters. At present, directions of application research are classified into three categories ; (a) Developments of high quality superconducting wires and tapes. (b) Developments of large bulk superconductors having high critical currents. (c) Developments of new superconducting electronic devices. All these application research must be based on how we can control those materials in quality and functions. In this report, the basic problems in research and development of high-Tc superconductors will be discussed and the prospect of future applications will be shown.
INTRODUCTION In the past ten years, more than twenty different kinds of high-Tc superconductors were discovered, and more than a half of them have critical temperature beyond 100K. The highest critical temperature is obtained in Hg-compound as high as 136K in ordinary atmosphere. This means that the operating temperature for future superconducting devices is widened from 4.2K to beyond 100K. On the other hand, the refrigeration technology has shown also a remarkable progress in its development, and this gives us a wide range of selections of the operating temperature. Generally speaking, the specific heats of solids depend on temperature strongly at low temperature regions, and the superconducting system operating at 77K has much larger heat capacity, sometimes more than 100 times, than that operating at 4.2K. This means that any systems or devices of high-Tc superconductivity are much more stable than that of low-Tc. Furthermore, the serious instability problem in the case of the low-Tc superconducting system, that is the quenching phenomena, may not be encountered in the high-Tc superconducting systems. These facts indicate that the market for the high-Tc superconducting systems must be quite wide enough for transportations, many industrial factories, electricity and electronics, and medicals. The high-Tc superconducting materials, however have several difficulties to reach real applications. This is due to the specific characters of high-Tc superconducting oxides. In this paper, it will be discussed how we have been overcoming those difficulties in high-Tc superconductors.
4
ICEC16/ICMC Proceedings
REQUIREMENTS ON MATERIALS FOR PRACTICAL APPLICATIONS Before discussing the requirements on materials for applications, we have to settle the operating temperature. From both, the economical and technological points of view in the refrigeration system, the higher operating temperature is better. However, the requirement to materials becomes severe in high temperature, since the characteristics of the material depend strongly on temperature. It seems to be plausible to chose 20K and 77K as the operating temperature. At 20K, the most of the characteristics of high-Tc superconducting oxides, Hg -, T1-, and Bi-based compounds and 1-2-3 compounds are, even at present, comparable or rather superior to those of low-Tc materials at 4.2K except the problem of thermal noise. Therefore, it is reasonable to settle the operating temperature for applications in the near future as 77K, and to discuss the requirements or targets of developments of high-Tc superconductors. Superconducting wires and tapes Taking into account the very wide future applications of superconducting wires and tapes, the requirements for critical current density must exceed 105 Amp / cm 2 at the magnetic field of 5T. This is rather moderate, if we consider the application to the nuclear fusion, however it covers still fairly a wide range of applications ; power cable, electric motor, maglev-cars and so on. At present, the silver-sheathed Bi-compound tape showed rather higher value in its critical current density at 20K, but the value is still lower at 5T, 77K. Bulk superconductors There are two kinds of applications of bulk superconductors ; one is using the magnetic levitation force in combination with ordinary permanent magnets, and the other is using the trapped magnetic field occurred by magnetic excitation. Both phenomena appear as the results of the pinning effect of quantized magnetic flux in the bulk. For consideration of the capability of the materials, the trapped magnetic field is a better measure. The trapped magnetic field is also strongly temperature dependent; it is roughly proportional to (Tc-To) 2, where To is the operating temperature. The requirement for the trapped magnetic field is 5T at 77K. It must be mentioned that the strength of the trapped magnetic field depends on the size of the bulk ; although it increases with the thickness of the bulk, it is assumed that here the thickness of the bulk is roughly 1 cm. The strength of the magnetic field at the surface of the ordinary permanent magnet is less than 1T, so it can be said that our requirement on the superconducting bulk is more than five times stronger than that of the ordinary permanent magnet. Superconducting electronic devices The requirements for the superconducting electronic devices are still not clear at present, since the image of future devices is not firmly established yet, eventhough many kinds of proposed ideas have been announced. But the basic needs for superconducting devices are very high speed operation and very low power consumption. In that sense, the single flux quantum (SFQ) device may be the most plausible candidate. This is simple extension of rf-or dc-SQUID and originally proposed by Likharev [1 ]. The simple logic devices were tried by Likharev [2], Okabe [3], Goto [4] and others by using Low-Tc materials (Nb) at 4.2K. This logic device is expected to have very high speed operation (10 psec) and very low power consumption (10pW). The logic devices using high-Tc materials have been tried recently by Westinghouse Co. [5], Hitachi Co. [6] and our laboratory [7]. The future image of the construction of the SFQ logic circuit is shown in Figure 1. In this figure, the superconducting single crystal is used as ground plane for the circuit and also as substrate for constructing the multilayer system, in which logic devices and wiring are included. In order to realize such a system, it is necessary to develop single crystal growth technology and high quality thin film technology.
ICEC 16/ICMC Proceedings PROBLEMS IN MATERIAL DEVELOPMENT In order to respond to the requirements of the future applications of high-Tc superconductor, we have to solve many problems in materials. The high-Tc superconducting oxides are composed of many different kinds of elements from three to five, and the crystal structures are also complicated. Then before discussing problems, several specific features of these material must be mentioned. Some of them are shown in Table 1 in comparison with low-Tc metallic superconductors. Passivation
Dimensionality Crystallinity
Metals
Oxides
Three Dim.
Two Dim.
not
sensitive
very
sensitive
Coherence Length Long ( > 1O0 A ) very short (< 30 A ) Boundary Effect
small
very large
Table 1. Superconductivity in Metals and HTSC
Insulater
Superconductor
Technologies (1)Single Crystal V///f//////////~f~///f////~/~/f~ Growth I .... ~-" IT_ ._-.., - ( 2 ) Homoepita xia I r/.l-././-j-./-.1-/-jj-./-/-./~/-~./.cjfj//////~jfjfj/~ Growth of High 12 ? '-Quality Thin Film (3)Multilayer Structure (4)Junction Fabrication (5)Passivation Basic
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All high-Tc superconductors, discovered until today, include so-called C u - O 2 planes in the crystal lattice, and the superconductivity occurs in these two-dimensional Cu-O 2 planes. Then, any lattice defect or impurity in the Cu-O 2 planes disturbs sensibly the superconductivity. Therefore, in any application, the Cu-O 2 planes in the crystal must be kept perfect as much as possible. Another specific feature is the short coherence length in these superconductors. It is estimated that the coherence length in a-b plane is around 2 0 ~ 3 0 A and less than 10 * along c-axis. This indicates that even very thin barriers among grains prevent the flow of supercurrent. These specific features suggest that the story of material problems must begin from the growth technology of pure single crystals of high-Tc superconductors. Single Crystal Growth Thin flakes of single crystal of YBCO and other 1-2-3 materials can be made by the flux method. But large single crystal of these materials can be grown only by the pulling method. Since those materials are incongruent, the special technology is necessary in growing large single crystals. A new method was invented in our laboratory in 1992 and it is called as the SRL-CP method. Single crystals of YBCO, NdBCO, SmBCO and PrBCO were grown by this method. Figure 2 shows a photograph of a typical single crystal of YBCO [8], and the cristallinity of this crystal investigated by X-ray diffraction reaches almost the same level as that of commercial single crystals of S r T i O 3 and MgO. It is rather surprising that the single crystal of YBCO reaches such grade of perfection nevertheless of its very complicated structure. It was also shown that these single crystals have also very high quality in electronic properties. We expect that we can obtain pure single crystals of 25 • 25 • 25mm 3 in the size in near future. Pinning Centers and Critical Currents The most important imperfections in superconductor are pinning centers, which pin the magnetic vortex. By this pinning effect, the critical current of superconductor increases. At present, it is considered that many kinds of imperfections can be pinning centers ; domain boundaries, dislocations, stacking faults and so on. However, it becomes clear that some kinds of the lattice defects having macroscopic size, from 100 .a. to several thousands A, survive as pinning centers at liquid N 2 temperature. This was confirmed in the YBCO bulk material containing finely dispersed particles of the Y2Ba~CuO5 phase [9]. The traces of ion bombardments in YBCO also increase the critical current.
6
ICEC16/ICMC Proceedings Pinning Center
Matrix ( V e r y Good Single Crystal)
lOmm
Figure 3 Image of "Pseudo Single Crystal"
Figure 2 Y123 single crystal grown along the c-axis
On the other hand, those materials having high critical currents showed rather good superconducting characteristics ; the critical temperature is almost the same as that of pure single crystals and the super-normal transition is also fairly sharp. These facts lead us to the new concept of pseudo-single crystal. The image of the pseudo-single crystal is shown in Figure 3. It is expected that the large pinning force comes from the boundary between the pinning center and the matrix, and the sharper the boundary, the larger the pining force. Probably this sharp change in crystal structure at the boundary (and also sharp change in order parameter) may be the origin of strong pinning force. In the field of applications, the magnetic field dependence of the pinning force is essentially important. Usually, the critical current decreases with increasing magnetic field and this prevents the wide range of applications of superconductivity. In 1993, Yoo. et al. in our laboratory [10] found that bulk of Nd~Ba2Cu30 v and Sm~Ba2Cu307, which were made by the OCMG (Oxygen Controlled Melt Growth) method, showed a prominent peak in magnetic field dependence of critical currents at 77K as shown in Figure 4, and a peak current density much higher than the Jc of the YBCO bulk in the high magnetic field region was observed. Further, they succeeded in increasing critical current of NdBCO in the lower magnetic field region (0 ~-2T) by introducing Nd4Ba2Cu20~0 particles. By those processes, it becomes possible to control the critical current in a wide magnetic field range ( 0 ~ 5 T ) at liquid N 2 temperature. They also found that the irreversibility field, which is the measure of pinning force of pinning centers, reached more than 10T at 77K. SOxlO 3
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OPERATIONAL PERFORMANCE During 1995 and the beginning of 1996 the cryogenic system has reliably provided the nominal cryogenic conditions necessary to perform experiments concerning the main systems of the LHC, operating for more than 1500 hours at 1.9 K and recovering from more than 50 resistive transitions. The temperature increase in the pressurised helium II bath was limited to 6 mK during nominal current ramping (10 A/s) and 50 mK during nominal discharge (-130 A/s), with steady-state temperature stability at 1.9 K better than a few mK, both in co-current and counter-current operation of the saturated helium II loop. In spite of the limitations reached in counter-current mode, the capability of cooling down the string from ambient temperature to 1.9 K in just more than a week has been maintained. The operational experience gained during the first run has allowed us to define control strategies such that all stand-by and transient operations can be performed unattended in fully automatic mode. In the same way the magnet powering does not require presence of skilled cryogenic operator. Opportunity was taken to perform other cryogenic test, such as the validation of basic instrumentation for LHC and the assessment of heat loads at different temperature levels.
COUNTER-CURRENT FLOW OF TWO-PHASE SUPERFLUID HELIUM Since the superfluid helium cooling loop is now fed by overflow from the phase separator (figure 2), acting as a weir at the upper end, the neighboring section of the corrugated heat exchanger tube sees countercurrent flow of liquid and vapor phases at saturation. As the heat load, and thus the mass flow-rates are gradually increased, the velocity of the vapor flow rises up to the point where it transfers sufficient momentum to block the gravity-driven, counter-current liquid flow, and thus prevents wetting of the heat exchanger tube. As the conditions of flow blocking are approached, the liquid level in the phase separator must sharply rise to compensate for the decrease in liquid velocity, and an instability develops. This phenomenon has been extensively studied in vertical or quasi-vertical flows of conventional boiling fluids, and is known as "flooding" in boiler tubes [4], but no experimental work had been done on superfluid helium in quasi-horizontal geometries.
ICEC16/ICMC Proceedings
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Fig.2 Superfluid helium cooling loop In order to assess limitations to the superfluid helium cooling loop brought by this effect, experiments were performed by increasing heat loads, while maintaining magnet temperatures stable, at different values of saturation temperature (i.e. vapor density) between 4.5 and 1.8 K. The onset of flow blockage can be clearly diagnosed from the sharp rise in the phase separator level, and the subsequent drift in the magnet temperatures (figure 3). The most critical conditions occur at low temperature, i.e. when the vapor density is lowest. On the 1.4 % slope, the 1.8 K liquid flow in the 4.15-cm inner-diameter corrugated tube is blocked at a vapor velocity of 6 m/s, corresponding to a heat load of 46 W, i.e. above the nominal demand of the LHC half-cells [3]. Although occurring at higher mass flow-rate between 4.5 and 2.2 K, blocking of the liquid flow in the heat exchanger tube may increase time for cool-down of the string from 4.5 K. Strategy for fast cool-down is to optimize the saturation conditions as a function of magnet temperature, in order to work with high-density, low velocity vapor, while preserving sufficient driving temperature difference. In this fashion, cool-down from 4.5 to 1.8 K could be achieved in 6 hours, only slightly longer than Joule-Thomson valve sizing could permit.
MAGNET RESISTIVE TRANSITIONS Coping with resistive transitions in superconducting magnets is a concern of main importance for the safe operation of the LHC. When occurring, it provokes the discharge of the magnetic energy stored in the coils (several MJ) into the static liquid helium bath, provoking rapid thermohydraulic transients At a current of 13.1 kA, corresponding to the LHC ultimate magnetic field of 9 T, rates of pressure increase up to 140 bar/s and pressure peaks reaching 12 bar are commonly observed. A fast responding protection system was designed at CERN to trigger electrically the opening of two quench relief valves after 120 ms, with full opening less than 170 ms after quench detection. During the latest experimental run, it was shown that one quench relief valve, located either on the lower part and low end of the cold mass (discharging liquid helium) or on its upper part and high end (discharging mainly in gaseous or supercritical phase) is sufficient to prevent the 1000 litres of the magnet helium inventory from reaching the design pressure of 20 bar. As shown in figure 4 the opening of the quench relief valve located on the upper part of the cold mass (the other being mechanically blocked) was delayed, in step of 10 ms, up to a total delay of 200 ms. In the dipoles (where the rate of pressure increase are the highest), the pressure peak increased by 2 bar with respect to no delay, still well below the 20 bar design pressure of the cold mass vessel. While keeping the maximum delay time for the electrical trigger, we lowered the pressure force on the valve seat down to 10 bar in order to force the valve to open on pressure. The pressure actuated opening gave a total delay time of 165 ms and a cold mass peak pressure of 13.5 bar. These tests were aimed at investigating the use of modified commercial valves in order to avoid the use of specially designed, fast-opening prototype valves as the ones in use [5]. The final tests were done using a standard and commercially available industrial valve, modified to fully open in less than 300 ms, mounted on the upper part and high end of the cold mass as one of the two fast opening valves. The test of this valve showed comparable discharge rates and pressure peaks in the cold mass.
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CONCLUSION The basic design technical choices for LHC and the modifications introduced after the first experimental tests have been validated, giving us confidence on the cryogenic system that is going to be built for LHC. The practical limitations of the counterflow system have been investigated and will give useful information for the definition of the final design which respect operational constraints. The tests performed on the prototype quench relief valves have given decisive information for the specification of the quench relief valves to be used in LHC. Analysis and simulation work is under way to understand the physics behind the process of energy transfer to the helium bath after a quench and its subsequent discharge and recovery. The string is now being prepared to assure 3000 hours of uninterrupted and fully automated operation at 1.9 K with continuous ramping and de-ramping to and from nominal current, to gain information on long-term behaviour of key components over the complete lifetime of the future machine.
ACKNOWLEDGEMENTS The authors wish to thank the LHC String Team, B. Gaillard-Grenadier and Th. Goiffon for the valuable work performed during the operation and maintenance of the LHC test string.
REFERENCES ~
2
e
o
5.
Evans, L.R., The Large Hadron Collider Project, paper presented at this conference B6zaguet, A., Casas-Cubillos, J., Flemsaeter, B., Gaillard-Grenadier, B., Goiffon, Th., Guinaudeau, H., Lebrun, Ph., Marquet, M., Serio, L., Suraci, A., Tavian, L. and van Weelderen, R., The superfluid helium cryogenic system for the LHC test string: design, construction and first operation, paper presented at CEC 1995, Columbus (1995) Lebrun, Ph., Superfluid helium cryogenics for the Large Hadron Collider Project, Cryogenics ICEC Supplement (1994) 34 1-8 Wallis, G.B., One-dimensional two-phase flow, McGraw-Hill (1969) Danielsson, H., Ferlin, G., Jenninger, B, Luguet, C., Milner, S.E. and Rieubland, J.M., Cryogenic performance of a superfluid helium relief valve for the LHC superconducting magnets, paper presented at CEC 1995, Columbus (1995)
Demands in Refrigeration Capacity for the Large Hadron Collider Ph. Lebrun, G. Riddone, L. Tavian and U. Wagner LHC Division, CERN, CH-1211 Geneva 23, Switzerland The capacity demands on the cryogenic installations for the Large Ha&on Collider (LHC) at CERN have been recently updated [ 1]. Unlike the LEP energy upgrade using superconducting acceleration cavities LHC will require high power refrigeration at 1.9 K, as well as non-isothermal cooling at 4.5 K to 20 K and at 50 K to 75 K. This paper presents the assessment of cryogenic capacity that has to be supplied by the eight refrigerators for LHC in relation with the foreseen operating modes of the machine. INTRODUCTION Based on today's knowledge the LHC collider will demand a total about 6.6 MW exergetic capacity. Half of this total capacity is to be supplied at 1.8 K for cooling the cold mass of the superconducting magnets. The rest is distributed mainly for non isothermal loads at 4.5 K to 20 K and 50 K to 75 K as well as for the cooling of current leads. Due to the exergetic losses which result from refrigeration at 1.8 K, the eight cryogenic plants will have a total equivalent capacity of 150 kW at 4.5 K. The reference design for the current leads is based on the use of high temperature superconductor (HTS) material leaving open the choice of cooling methods [2]. The precise assessment of the heat loads is an important task during the planning phase of the project and has to be kept under fight control. TEMPERATURE LEVELS The staging of temperature levels envisaged for the LHC cryogenic system are" - thermal shielding between 50 K and 75 K as a first major heat intercept, sheltering the cold mass from the bulk of heat inleaks from ambient; -- non-isothermal cooling of the beam screens between 4.5 K and 20 K which will shield the cold mass from synchrotron radiation and image currents produced by the high-energy, high-intensity hadron beams; quasi-isothermal superfluid helium cooling the magnet cold mass at a maximum temperature of 1.9 K; isothermal helium cooling special superconducting magnets in insertion regions, superconducting acceleration cavities, and the lower sections of HTS current leads at a saturation temperature between 4~ K and 4.7 K, gaseous helium cooling the resistive upper sections of HTS current leads, in forced flow between 20 K and ambient.
-
-
-
STATIC AND DYNAMIC LOADS The term "static loads" is used for all heat loads which are not dependent on the machine operation. This comprises the heat inleaks due to conduction and radiation to the low temperature part inside the cryostats plus the thermal load by pure conduction in all feedthroughs. The term "dynamic loads" is used for all heat loads which are dependent on the machine operation. This includes loads induced either by the excitation current of the superconducting magnets i.e. in current leads or residual resistive areas of coil splices, or by the circulating particle beams i.e. synchrotron radiation, resistive heating due to image currents or geometrical singularities in the beam channels and 95
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particle losses. These loads depend strongly on the energy and intensity of the circulating beams and are estimated for the nominal and ultimate operating conditions defined below. MODES OF OPERATION The modes of interest for the cryogenic system are: - "nominal operation" at 7 TeV beam energy, 2 x 0.536 A beam current and 1 x 1 0 34 c m "2 s "1 luminosity; - "ultimate operation" at 7 TeV beam energy, 2 x 0.848 A beam current and 2.5 x 1034 c m 2 s 1 luminosity; "low beam intensity operation", still with full beam energy and thus full excitation current in the magnets but negligible beam current; "injection standby", characterised by negligible resistive dissipation and beam-induced heat loads in the magnets.
-
-
THE ARCS The eight arcs, which form the regular part of the LHC will be composed of 46 half-cells, each consisting of three main dipoles, one main quadrupole plus a number of correction magnets [1]. The expected static loads of the main elements in the half cells has been assessed with good confidence and experimentally verified through several independent, yet converging approaches [3-6]. Parallel to the magnet cryostats in the arc runs the cryogenic distribution line from which each halfcell is supplied with cryogenic fluids. The heat loads of this line are estimated based on conceptual design drawings and will have to be validated in future prototype tests. Table 1 shows the total heat load budgeted for an LHC arc in the four machine operation modes described above. Table 1 Total heat load for an LHC arc (no contingency included) Operating mode
Nominal operation Ultimate operation Low-beam-intensity Injection standby
Temperature levels 50- 75 K 4.5 - 20 K 1.9 K
[w]
[w]
[w]
15800 15800 15800 15700
1930 3630 310 310
1070 1040 920 660
Comments
static & dynamic nominal static & dynamic ultimate static & magnet current induced pure static load
THE INSERTIONS The insertions consist of 0.5 km long sections connecting the arcs. They include dispersion suppressor, beam separation and recombination, RF acceleration and beam focusing sections on each side of the collision points. Each cryogenic sector includes two half insertions at the both ends of an arc. Since the detailed design of the superconducting magnets in the insertions has not yet been finalised, their expected heat loads have been assessed by approximate scaling from the better defined arc magnets. The results, displayed in Table 2, reflect the variations in geometry, layout, functions and operating modes of the different LHC insertions. They also include specific components- for example, superconducting acceleration cavities, which require significant amounts of cryogenic refrigeration as well as reference magnets for field tracking measurements. Also appearing in these tables are the localised heat loads in the current leads of the different excitation circuits, and based on the use of high-temperature superconductors. Special heat loads have to be accounted for all elements close to the high luminosity physics experiments due to absorption of secondary particles in the cold mass of the superconducting magnets located close to the experimental areas. The values for the whole machine are" - at the 1.9 K level, from 500 W (nominal) to 1250 W (ultimate) in the inner triplets of low-beta quadrupoles, as well as from 54 W (nominal) to 134 W (ultimate) in the dispersion suppressor magnets;
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- at the 4.7 K level, from 240 W (nominal) to 600 W (ultimate) in the beam separation and recombination dipoles. Table 2 Total heat loads for the half insertions of the LHC machine at nominal operating conditions (no contingency included) Temperature Levels 5 0 - 75 K
4.5 - 20 K
4.7 K
1.9 K
Gas helium consumption 2 0 - 300 K
1300 1350 800 1250 1150
125 130 120 130 120
100 50 5 380 5
300 150 80 100 120
11 13 2 13 12
Half Insertion
[w]
High-luminosity insertion Low-luminosity insertion Beam-cleaning insertion RE insertion Beam-dump insertion
[w]
[w]
[w]
[g/s]
DEMANDS ON THE SECTOR CRYOPLANTS The LHC sectors stretch over 3.3 km between the insertion points. The cryogenic load for each sector comprises those a full arc and two half insertions at each end. A special load to be added for each sector is that resulting from a continuous loss of particles escaping the collimation sections, which may result in a locally deposited heat load estimated at 55 W (nominal) and 92 W (ultimate), over a length of a few tens of metres corresponding to the region of aperture restriction. The total installed cryogenic power to cope with the expected loads for the sectors is determined using a contingency calculation to include for uncertainties and overcapacity. For the heat loads at the 50 75 K and the 20 to 300 K level: Qinstalled = Max [ Fov" Qnominal ; Qultimate]; for the heat loads at the 4.5 - 20 K, 4.7 K and 1.9 K level: Qlnstalled = Max [ Fov" (Fun" Qstatic + Qdynamic nominal) ; Fun" Qstatic + Qdynamic ultimate ].
As factor for overcapacity, Fov, 1.5 is used, for uncertainty, Fun, 1.25. Table 3 shows the installed cryogenic power which results for the two types of sectors of the machine. About 250 W of the load at 1.9 K listed in Table 3 result from superheating the helium vapour in the 3.3 km long cryogenic distribution line and do not contribute to the cold compressor flow. Table 3 Installed cryogenic power requirements of the LHC Temperature Levels
Sector 5 0 - 75 K
[kW]
High-load sector Low-load sector
31.0 30.0
4.5 -20 K
[kW] 4.30 4.30
4.7 K
[kW] 0.80 0.65
1.9 K
Gas helium consumption 20 - 30O K
[kW]
[g/sl
2.80 2.45
35 23
DAILY LOAD VARIATIONS The operation phases which the machine will undergo during a 24 hour period will be four hours of injection standby, during which the storage ring will be adjusted and filled with particles, half an hour of current ramping to reach the operating energy, 19 hours of operation at nominal or eventually ultimate conditions, half an hour of current ramp down to prepare for the new injection. Current ramping will generate an additional heat load due to resistive dissipation in the superconducting magnet splices.
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Figure 1 shows the development of the load on the 1.9 K level during the injection and ramp phase leading from the "injection stand-by" to the "nominal" operation mode. Although the rapid changes will be partially buffered by the heat capacity of the pressurised helium II bath, the variations in demand require load adaptation of the cryogenic system over a large dynamic range. 160
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CONCLUSIONS The demands of refrigeration capacity for the LHC machine have been assessed with increasing accuracy during the recent years. As the project definition is finalised further updates of the heat loads will be necessary. Special attention has to be given to all components which at the moment are still uncertain in their performance like the insertion magnets or assessed by theoretical calculations like the HTS current leads. REFERENCES The LHC study group, The Largo Hadron Collider, Conceptual Design CERN/AC/95-05(LHC) Ballarino, A., Ijspeert, A. and Wagner, U., Potential of high-temperature superconducter current leads for LHC cryogenics, paper presented in this conference Jenny, B., Cameron, W., Riddone, G., Rohmig P. and van Weelderen, g., Design and construction of a prototype superfluid helium cryostat for the short straight sections of the CERN Large Hadron Collider, Adv. Cryo. Eng. (1994) 3 9A 663-670 Benda, V., Dufay, L., Ferlin, G., Lebmn, Ph., Rioubland, J.M., Riddone, G., Szeless, B., Tavian L. and Williams, L.R., Measurement and analysis of thermal performance of LHC prototype dipole cryostats paper presented at CEC'95 Columbus (1995) B6zaguet, A., Casas-Cubillos, J., Flemsaeter, B., Gaillard-Grenadier, B., Goiffon, Th., Guinaudeau, H., Lebrun, Ph., Marquet, M., Serio, L., Suraci, A., Tavian L. and van Weelderen, R., The superfluid helium cryogenic system for the LHC Test String: design, construction and first operation paper presented at CEC'95 Columbus (1995) Dufay, L., Ferlin, G., Lebrun, Ph., Riddone, G., Rieubland, J.M., RijUart, A., Szeless B. and Williams, L.R., A full-scale thermal model of a prototype dipole cryomagnet for the CERN LHC project Cryogenics 34 ICEC Supplement, (1994),. 693-696
Simulation Program for Cryogenic Plants at CERN
E. Melaaen, G. Owren*, A. Wadahl and U. Wagner LHC-Division, CERN, CH-1211 Geneva 23, Switzerland *Department of Refrigeration and Air Conditioning, NTNU, 7034 Trondheim, Norway
A steady-state simulation program for helium refrigeration plants at CERN has been developed. The objective was to have a tool available for analysing existing plants in connection with the LEP project, but also for studying system solutions for future plants in the LHC project. The program can simulate from single components up to complete cycles, but it can also be used for closer study of heat exchangers and expanders in a specified process. Each component is modelled by writing its energy balances, the equations are put into a unique matrix and solved. Since refrigeration cycles are mainly closed-loop processes, the program solves the complete equation system simultaneously.
INTRODUCTION To reach the necessary beam energy in the Large Electron Positron collider (LEP2) at CERN, superconducting cavities operating at 4.5 K are installed. Four 12 kW @ 4.5 K helium cryogenic plants, supplied by LINDE and L'AIR LIQUIDE, are installed for the task [1 ]. There are also other cryoplants at CERN used in connection with test facilities. During his stay at CERN, G. Owren started to develop a tool (CryoSim) for analyses and studies of existing refrigeration plants. This is the continuation and development of his work. The tool is now used to simulate existing plants, but also for process evaluation of future cryoplants for the Large Hadron Collider (LHC) project. SCOPE AND CAPABILITIES OF CRYOSIM CryoSim is a robust steady-state simulation program developed for cryogenic plants, but can also be used for process simulations in general. Thermophysical properties for helium given in the commercially available package HEPAK [2] are installed. Data from HEPAK are based on fundamental state equations. Since only helium data are implemented, the program has limitations for an extended use. Other process fluids can be installed in the future. CryoSim is especially developed for off-design calculations of existing plants. The program is a useful tool for process calculations and optimization of helium liquefiers, and has been helpful in understanding and assessing the performance and efficiency for various cryoplants. Other tasks for the program can be thermodynamics studies and as a tool during specification of cryoplants. The first version of CryoSim had some limitations since the turbine efficiencies (11) were taken constant and independent of the operation conditions. In the same way the heat-transfer impedance in the heat exchangers (UA) did not vary with the mass flow. This is improved, and new models for turbine efficiency and heat-transfer coefficient varying with the flow rate are implemented. The program can now be used for simulating situations with varying cooling power or at off-design conditions with better results. 99
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For a given process the flow-scheme has to be described in an input file following a predefined structure. The input file is mainly divided into two parts. The first part describes each component such as heat exchanger, valve, etc., and all necessary characterizations of the component such as pressure drops and efficiencies are given here. In addition inlet and outlet streams for each component and the connections between all components are specified. The second part of the input file gives the information about all the mass streams in the process. To specify a stream, flow rate, pressure together with enthalpy or temperature have to be given. From the given data in the input file, the program understands if a specified stream represents a boundary or an initial condition. The following components are modelled and included in the program; boiler, brake, compressor, expander, flash drum, heater, heat exchanger, mixer, pipe, splitter, subcooler and valve. The principle for modelling each component is the same, and the structure of the program makes it easy to install new component models. Heat exchangers are modelled in two different ways; one is based on a LMTDmethod (called hx), while the second method is more rigorous (called rhx). The rigorous method divides the heat exchanger into a specified number of blocks, and each heat-exchanger pass in a block is modelled separately. With this method it is possible to get out a thermodynamic description of the condition in the heat exchanger (e.g. temperature profile). Also a control facility is available in the program which can be used to update boundary conditions until specified set-points (e.g. temperature or pressure) are reached. The program can calculate the flow sheet specified in the input file in two different modes; simulation or design. In the simulation mode the process is solved as a steady-state simulation from boundary conditions and specified component parameters. Design mode uses the specified stream data to calculate the heat-transfer coefficient for heat exchangers and isentropic efficiency for expanders. The calculated results are stored in a specific output file. This file contains flow sheet data, balances for mass and energy, exergy analyses, stream data (temperature, enthalpy, entropy, density, pressure, flow rate and quality), component data (pressure drop, UA, efficiency e.g.), calculation time and information about iterations. Temperature profiles, pinch and NTU for heat exchangers are also available. To work with CryoSim it is necessary to have some basic knowledge about UNIX | and one editor available at UNIX | (e.g. emacs). For the moment the program is only running on UNIX | Figure 1 shows an example of one process together with part of its input file. The process can be regarded as the lower end of the cold box for one of the 12 kW plants at CERN. The first line in the input file describes the main cluster of components here called EXAMPLE. The second line indicates if the calculation mode is design or simulation. After this line starts a list of all member components in the cluster EXAMPLE. All components with their characterization are written in a general and fixed structure. The first line for a component starts with an identification of the component type (such as expander and rhx) and an identification name. The information describing a component is finished by the word end. The example shows the structure of one heat exchanger and one expander model. The last part of the input file specifies information of each stream in the process. Every line starting with mass gives either boundary or initial condition for one stream. PROGRAM STRUCTURE CryoSim is based on an object-oriented technique and written in the program language ANSI C | which gives the program large flexibility. The equation system for each component is set up in a unique matrix coveting the equations for the whole plant. Since refrigeration cycles mainly are closed-loop processes, the program solves the total number of equations as one large system (simultaneously) with a type of Gauss-Jordan elimination method. The results is a robust program. The components are modelled by the energy balances; therefore, the enthalpy for each stream is solved. The mass flows and pressures are updated separately by following the streams from boundary conditions and further through all the streams in the flow-scheme. For stream splitters and flash drums either split factor or draw-off rate have to be specified in the input file. Similarly must the pressure drop inside each component or outlet pressure for
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a component be specified in the input file. For a heat exchanger an option will later be available for calculating the pressure drop dependent on the flow rate.
compound EXAMPLE mode simulation rhx HE1 feeds mass str l mass str5 prods mass str2 mass str6 paraml 101-1 pspec 1 0.1 0.05 design 0 38000.0 design 1 38000.0 end expander T 1 feeds mass str2 prods mass str3 work wb 1 pspec 1 4.5 param 1 0.7 end
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Figure 1 Shows the principle for an input file describing a given process [3] The equations based on the energy balances are put into the matrix and solved. A matrix solved by the Gauss-Jordan method normally requires constant coefficients, but a matrix system for a thermodynamic process has non-constant coefficients. An iteration procedure is therefore necessary for solving the system. For each iteration the coefficients are kept constant and the matrix is solved. The coefficients are updated between each iteration. When the solution vector has not changed between two iterations, the solution is reached. Figure 2 shows the program structure. Each box represents one directory, and subroutines in one directory are compiled in a library. All libraries are compiled together in the directory steady, and the result is the executable file sire. Directory CryoSim is organized into three main parts. The mat library represents the mathematical library with equation solvers, while cryodata and hepak contains the libraries for thermophysical properties. Basic contains the main structure of the program and describes the organization and storing of data. Two important structures defined in basic are the organization of streams and components. The two sub directory steady and dynsim contain respectively a steady-state
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and a transient simulation program, dynsim is under preparation and, therefore, not discussed in this paper. In the directory steady steady-state models of the components are stored, and here is the main program compiled to the executable file sire. To run the program, sire must be available for the user. Plot is meant to be a tool for graphical presentation of calculated results as temperature profiles inside heat exchangers and TS-diagram for a calculated process. Plot is still under development. i
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executable program sire Figure 2 Shows the directory structure of the program and links to other applications [3] CryoSim is partly tested out at CERN by comparing simulated results and measured data from running plants. These studies show good validity for the program, but a closer study is required and will be done in the near future. Especially important is it to test the heat-exchanger models and expander models at design and off-design situation. A thorough description of the program CryoSim and documentation of the validity of the models are given in [3-5]. CONCLUSION CryoSim is a robust calculation program for helium processes and cryoplants, which can be used for simulation and for limited design cases. The program is useful during analyses of existing plant and for evaluation of new processes. Since the total equation system is solved simultaneously, the program is adjusted to solve closed-loop processes as refrigeration cycles. The object-oriented structure gives the program large flexibility and possibilities for further development. CryoSim is tested at CERN, and preliminary results show good validity [3-5]. REFERENCES G0sewell, D. et al., Cryogenics for the LEP200 Superconducting Cavities at CERN, Particle Accelerator Conference (PAC93), Washington D.C., 1993 Cryodata inc., User's Guide to HEPAK, Version 3.21, Colorado, January 1993 H~ye, G.K., A Simulation Program for Helium Cryoplants, Thesis, Norwegian Institute of Technology, Trondheim, 1994 Wadahl, A., CryoPlant Process Simulation Program, Thesis, Norwegian Institute of Technology, Trondheim, 1995 Melaaen, E., User's Guide to CryoSim and System Documentation for CryoSim, in preparation, CERN, Geneva, 1996
Operation of the Cryogenic System for Superconducting Cavities in LEP
M. Barranco-Luque, S. Claudet, Ph. Gayet, N. Solheim, and G. Winkler LHC Division, CERN, CH-1211 Geneva 23, Switzerland
At CERN the upgrade of the LEP e+e - collider towards higher beam energies is under way by installing superconducting cavities in the ring. In 1995 superconducting cavity modules have been operated together with ambient temperature copper accelerating cavities allowing for a first step of energy increase. We report on the experience with the operation of the LEP cryogenic system. Particular attention is given to stability, automatic control, and reliability. Failure analysis and redundancy programs are presented which should further increase the availability of the cryogenic system in the environment of a large high energy particle collider.
INTRODUCTION The upgrade of the LEP e+e - collider from 45 GeV to 96 GeV per beam is under way by installing gradually up to 272 superconducting cavities on both sides of the four interaction points in the ring. These 352 MHz cavities are assembled in 4-cavity modules and cooled by four large cryoplants with 12kW equivalent capacity at 4.5 K. In 1995 a total of 16 modules have been operated at three LEP points in addition to the ambient temperature copper accelerating cavities allowing for a first step of energy increase from 45 to 70 GeV. After the LEP winter shut-down 1995/96 all four 12 kW/4.5 K refrigerators will be in operation with 35 out of the final 68 cavity modules, the totality of which will be operational in 1998. THE LEP2 CRYOGENIC SYSTEM The cryogenic system at each of the four interaction points of LEP, described in [1] and earlier publications, consists of a cryoplant [2,3,4] with an equivalent cooling capacity of 12 kW at 4.5 K and its associated liquid helium distribution system, a pair of about 200 m-long supply-and-return transfer lines to feed the 8 or 9 superconducting (sc) acceleration cavity modules on each side of the point. First experience with the 12 kW plants was reported in [5,6]. A description of the liquid and gaseous helium circuits inside the 11 m long 4-cavity modules, which are treated as independent cryogenic units for cooling and controls, is given in [7]. In addition to the sc cavities the low-beta quadrupoles at two LEP points are also being cooled by the 12 kW plants. As an example of the layout the cryogenic system of LEP point 2 is shown in Figure 1. CONTROLS AND AUTOMATIC OPERATION The industrial process control system [8], purchased by CERN apart from the cryoplants, has to handle equipment at 4 points evenly-spaced around the 27 km long LEP ring, with a central control room in the office building of the operation team. At each point a local control room is installed and 5 programmable control units are distributed on the surface and under ground. 1800 input/output channels are connected to the process of each cryoplant and the associated modules. The control system is programmed for fully automatic operation including cool-down, warm-up, restart 103
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Transfer lines connecting upper and lower cold box
12 KW Upper Cold Box ( 300- 20K ) j
Machine tunnel RA 27
Helium
Service Tunnel UA 27 Transfer lines terminal by-pass Suoerconductino cavitv module LHe distribution system in machine tunnel RA 27
Low beta Superconducting quadrupoles 12 KW Lower Cold Box ( 20 - 4.5 K ) in US 25
LHe distribution system for 32 sc cavities in machine tunnel RA 23
Figure 1 after utility failure and adaptation of the plant capacity to reduced load. Operation can be efficiently supervised from the central control room by a small operator team, on duty only during normal working hours or on automatic call in case of failure or alarm on sensitive parameters. OPERATING CONDITIONS FOR CAVITY MODULES The inlet and outlet valves of each module are controlling the level and the pressure of the liquid helium bath. In addition a compensation of the induced dynamic radio frequency (RF) load is implemented by means of electrical heaters. An algorithm is calculating the necessary compensation power directly from a RF field strength signal and the pre-set cavity quality factor. It is important to note that the outlet valve of the module throttles the flow and thus attenuates the influence of pressure variations from one module to its neighbours, as well as variations possibly induced by the compressors suction pressure. Table 1 presents the operating conditions achieved during steady state operation and when a RF field step with typically 400W load change is applied. Table 1 Cavity module operation conditions and stability
Pressure Level
Nominal conditions 1250 mbar 800 mm
Steady state variations +/- 2 mbar +/- 5 mm
Variations by RF steps +/- l0 mbar +/- 10mm
RELIABILITY AND AVAILABILITY The four 12 kW plants have now accumulated a total of 34000 hours of operation. Fault statistics for the complete cryosystem operation (including control system) during last years' LEP runs are presented in Table 2. This includes the final 12 kW plants, as well as the now replaced 6 kW plant. It turns out that interruptions of operation were mostly related to utility failures (electricity, cooling water, etc.) and not to specific cryogenic problems. As a consequence of the failures in 1995, the non-availability of the cryosystem for RF operation of the cavities was 32 h, including normal recooling and refilling time, or 0.4 % of the total LEP running time.
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Table 2 Cryosystem fault statistics Year
Number of installed modules 1992 2 1993 3 1 1994 4 3 .....i995 ......................................4................................ 8 4 .
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LEP point 2 2 6 2 6 2 6 8 .
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Cryoplant type & operation time during LEP run 6 kW 6200 h 6 kW 3800 h 12 kW 1400 h 6 kW 4600 h 12 kW 1500 h 12 kW -3900 h 12 kW 3900 h 12 kW 840 h .
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Cryoplant stops cryo faults-total faults 1-7 2-8 0-2 3-14 1-3 0-3 0-4 0-0 .
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After plant stops the cryogenic system will introduce unavoidable delays to the RF operation during reestablishing of steady state conditions. Table 3 shows the recovery times experienced during the last years with a small number of modules (4 to 6) and the extrapolation for 16 modules connected to one plant. It also indicates the characteristic times of the cryosystem for the most relevant operation modes. Table 3 Characteristic times for the cryosystem
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CONSOLIDATION TASKS Lame storage tanks In addition to the present 10 helium gas storage tanks (each 75 m 3, 20 bar, vertical axis) at each of the four LEP points, 3 large tanks (each 250 m 3, 20 bar, horizontal axis) will be installed to allow for one complete refill of the modules in case of accidental loss of helium. Redundanev and maintenance For cost reasons the plants where originally specified and built without redundancy for critical items. To ensure minimum downtime most spares are now on stock. In view of the rather long intervention time for compressor repairs and future needs of increased flowrates for plant upgrades for the next accelerator project LHC, fully equipped redundancy compressors, one for each of the two pressure stages of each plant, are being procured from industry. It is also planned to install redundancy for vital components of the cooling water system, similar to those already in place for the compressed air supply. Most crucial for the reliability of components is the future preparation of detailed maintenance plans and the thoroughly executed preventive maintenance during the winter shut-downs of LEP. H~lium management An operating 12 kW plant is filled with 2500 Nm3 of helium. During installation, commissioning and testing about 4 times this amount has been used for each plant. The total contents of the running v
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cryosystem at one point with 18 modules (550 Nm3 each) will be 15000 Nm3, 10% of this amount were lost last year during maintenance or module installation, 15% for leaks and incidents with incomplete helium recovery. Efforts must be spent on reducing these losses during next years. Ootimisation of ooeration mode~ Further work is planned to optimise the operation modes, in particular the recooling and refilling, to reduce restarting delays of the increasing number of modules. A close follow-up of the dynamic thermal losses in the modules will be implemented to minimise unnecessary load induced by the RF compensation heaters. Organi sation of operation CERN has established a contract with an experienced company to ensure operation and maintenance of all cryogenic installations at CERN. The intention is that this company gradually during a three years period will take over the operation and maintenance responsibilities with guarantee of results. While this outsourcing policy, which is now also applied to operation tasks on complex technical equipment, gives promising results after the first 9 months of the contract, much further training of new personnel and transfer of specific experience will still be necessary in the coming years. CONCLUSIONS AND OUTLOOK Operating experience with the LEP2 cryogenic system, which is at present the most powerful and efficient helium system world-wide, is very encouraging. After replacement of a defective heat exchanger in the upper coldbox of one plant, roughly the same cooling capacities have been reached at all four points. Reliability of the cryoplant components and of the control system turned out to be remarkably good. Some further work on automatic procedures, redundancies and preventive maintenance, should allow to face successfully the future demands for full capacity and high availability of the cryogenic system as part of the upgraded LEP collider. REFERENCES Gfisewell, D., .Barranco-Luque, M., Claudet, S., Erdt, W., Frandsen, P.K., Gayet, Ph., Schmid, J,, Solheim, N.O., Titcomb, C. and Winkler, G., Cryogenics for the LEP200 Superconducting Cavities at CERN, Proc. of thr P~rticlr Accel. Conf, 1993, 2956-2958 Chromec, B., .Erdt,W.K., Gfisewell ,D., LOhlein ,K., .Meier, A., .Senn, A.E., Saugy, T., Solheim, N.O., Wagner, U., Winkler, G., Ziegler, B., et. al, A High Efficient 12 kW Refrigerator for the LEP200 Project at CERN, Proc. of the IISSC 1993. Supercollider 5. Plenum Press (1994) 95-100 Gistau, G. and Veaux, J., A 12/18 kW at 4.5 K Helium refrigerator for CERN's LEP superconducting accelerating cavities, Cry_ogenics (1994) 34 ICEC Supplement 103-106 Claudet, S., Erdt, W., Frandsen, P-K., Gayet, P., Solheim, N.O., Titcomb, C. and Winkler, G., Four 12 kW/4.5 K Cryoplants at CERN, Cry_ogenics (1994) 34 ICEC Supplement 99-102 Winkler, G., Gayet, Ph., Gtisewell, D. and Titcomb, C., Cryogenics Operation and On Line Measurements of RF Losses in the SC Cavities of LEP2, Particle Accel, t~0nf, 1995 Dallas Barranco-Luque, M., Claudet, S., Dauvergne, J.P., Erdt, W., Frandsen, P-K., Gayet, P., Gtisewell, D., Lebrun, Ph., Schmid, J., Solheim, N-O., Titcomb, C., Wagner, U. and Winkler, G., Conclusions from Procuring, Installing, and Commissioning six large-scale Helium Refrigerator at CERN, CEC/ICMC 1995 Columbus Barranco-Luque, M. and Gfisewell, D., Thermal Loss Analysis of Cryostats and Accessories for the Superconducting Cavities of the LEP Energy Upgrade, Proc.of the European Particle Conf. (1994) 2455-2457 Gayet, Ph., Claudet, S., Frandsen, P-K., Juillerat, A., Kuhn, H.K., Solheim, N.O., Titcomb, Ch., Winkler, G., Wolles, J.C. and Vergult, P., "Architecture of the LEP2 Cryogenics Control System' Conception, Status, and Evaluation", Cry_ogenics (1994) 34 ICEC Supplement 83-86
Thermodynamic Booster for the CERN Omega Cryoplant
F. Haug, J.-P. Dauvergne, H. Rieder, P. Chaffard CERN - LHC Division, CH-1211 Geneva 23, Switzerland
The 800 W @ 4.5 K cryoplant, commissioned in 1971 for the superconducting magnet of the OMEGA particle detector, uses a large reciprocating compressor designed for the delivery of helium at a rate of 216 g/s at 20 bar. A recurrent problem, which overshadowed the good performance of the plant, was the rapid wear of the labyrinth-type, compressor piston seals. This was due to the high temperature of the compressed gas, a consequence of the rather high compression ratio per stage. Even very frequent maintenance did not prevent progressive loss of efficiency and lack of cooling power. A costly replacement of the compressor, by a better suited modern machine, was foreseen. Eventually, however, the problem was solved in a much cheaper way by inserting an auxiliary cooling stage between the compressor and the cold box. This was done with a commercial, fully automated, R22 refrigerator which reduced the helium temperature, at compressor suction and delivery, from ambient to about 10~ The result was not only an increase in cooling power of nearly 10 %, but also a reduction in the labyrinth wear to a level hardly expected - even by optimists!
INTRODUCTION The first large superconducting magnet installed at CERN was for the OMEGA multipurpose particle detector which permits experiments in a magnetic space of about 1.5 m height and 3 m diameter with a central field of 1.8 T. The cryoplant for its cooling was supplied by Sulzer, Switzerland and has been in operation for 25 years since its commissioning in 1971. The rapid wear of the labyrinth-type pistons seals of the compressor became a recurrent problem. This led to deficiencies in the cooling performance of the refrigerator such that it was barely sufficient to cool the magnet. After a brief description of the OMEGA refrigeration cycle, the problems encountered are analysed and the technical approach of "thermodynamic boosting" is explained. The flow scheme of the cycles are presented. The system has been successfully in operation for more than five years. THE OMEGA MAGNET AND THE CRYOGENIC SYSTEM The OMEGA magnet is made of two coils with a 3m inner diameter. They are mounted in a horizontal plane 1.5 m apart on a vertical axis. The coil is split into six, identical, double pancakes, all electrically connected in series. Each coil is powered via a set of two current leads at 4800 A. The coils are wound with a hollow superconductor, 18 x 18 mm 2, which has a central, rectangular, cooling hole of 9 x 9 mm 2. The refrigerator was designed for 500 W @ 4.5 K, 2.5 g/s for the current lead cooling, and 4000 W at 77 K for shield cooling (power is withdrawn from the helium cycle to a LN2 circuit via a He/N2 heat exchanger). The liquid nitrogen is circulated with a rotary pump. The forced flow cooling at 4.5 K is done with a J.T. flow of 45 g/s of supercritical, subcooled, helium. This at a pressure of 10 bar at the inlet to the coil system and 5 bar at its outlet. The flow is split into four branches i.e. two per coil and is subcooled three times in a cryostat by boiling helium at 1.25 bar each time before flowing through an individual section of the coil. Reference is made to publication [ 1]. An additional particularity, which adds to the complexity of the refrigerator process, is the "artificial" increase of delivery mass flow of the piston compressor. This is done by reinjection of part of the total flow at the exhaust of the first turbine back to the suction side of the third compression stage permitting to adjust and maintain the mass flow of the first turbine. The delivery mass flow is, therefore, in this case higher than the suction mass flow. The compressor is a labyrinth sealed, 3-stage, 4-piston machine, where two pistons form the first stage to provide the rather large mass flow of around 215 g/s (design value!). The helium gas is chilled after each compression stage - at interstage and after coolers with water from cooling towers to ambient temperatures. 107
108
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Figure 1 Maintenance of labyrinth sealed piston compressor TECHNICAL SOLUTION Our thermodynamic boosting system was designed to help overcome the problems described above. The intentions were to: reduce wear and compression temperature, increase the mass flow, increase the thermodynamic efficiency of the refrigerator cycle, ease the of operation of the plant and reduce the maintenance costs to a minimum. Two helium/Glycol-water heat exchangers were inserted into the helium circuits between the compressor and the cold box (see schematic of fig.2). Both are of longitudinal tube type with semicounterflow of the cooling agent Glycol-water. The heat exchanger at the 1.0 bar suction level, named "WT2", with length 1.8 m and 325 mm outer diameter, consists of 189 parallel tubes each of 16 mm diameter and 1 mm wall thickness, with an active core length of 1500 mm. The heat exchanger at the 20 bar delivery side contains 48 tubes of the same type and an active core length of 1800 mm. Both provide a cooling capacity of around 20 kW with a heat exchange coefficient, between helium gas and cooling agent, of 142 W/m2K for the low pressure side and 400 W/m2K for the high pressure side. In order to minimise additional entropy production in the helium flow the pressure drops for both heat exchangers were specified to be below a constringent value of 1% of the absolute pressure. For the suction side the pressure drop was observed to be 8 mbar and for the delivery side it was even lower.
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PROBLEMS AND THEIR ANALYSIS Despite the sophisticated, thermodynamic approach, to refrigeration requirements in the "early days" of large scale cryogenics, it was observed over the years of operation that the performance reserve of the refrigerator was very narrow and was continuously decreasing. This was particularly apparent in the warm seasons when ambient temperatures went above 25~ causing a parallel temperature rise in the cooling water for the compressor. This water being chilled by a cooling tower. As a consequence, the helium temperature at the suction side of the compressor and at the other two compression stages increased. Hence the mass flow decreased and the compression temperature, after each stage, increased. This at a constant pressure ratio caused progressively faster wear of the compressor labyrinth piston seals and the cylinder walls. Also the performance of the refrigerator decreased. In order to avoid a break down of the cryoplant an upper limit of compression temperature had to be respected. To achieve this the pressure ratio was reduced by the operators to obtain a delivery pressures of 18.5 bar and less instead of 20 bar. The wear of these seals produced another negative effect which was the reduced sealing efficiency of the labyrinth pistons. This led to a gas flow from the compression side to the suction side of the double acting pistons which added to the wake space of the remaining, not expulsed gas in the cylinder. The decreasing compression efficiency could be measured again by an increase of temperature. Thermodynamically this can be expressed in terms of the polytropic exponent "n" which increases. Following costly maintenance work, data was finally taken in the 1990's, during the running of the compressor, and the situation was then analysed. We have observed that with 6 months of operation, after maintenance of the compressors, the following: a) a loss in mass flow of 6.5 % occurred, b) increase in the polytropic exponent n (decrease of compression efficiency) by 7.0 % for the first stage, 8.6 % for the second and 7.5 % for the third stage, leading to increased temperatures. PERFORMANCE LIMITS The negative changes of the cryogenic performances were perceptible: 1) At the level of the mass flow rate which decreased sometimes as low as 86 % of the design value to 186 g/s during critical periods. 2) By the performance reserve of the refrigerator, at the 4.5K level, which in a normal case should be situated above 100 W reached 0 Watt. 3) By the liquid nitrogen cooling agent for magnet screens and mechanical supports which could not be maintained at the normal value of 76K by the helium/LN2 heat exchanger and approached temperatures of close to 80K. This endangered the nitrogen turbo pump which may fail by cavitation of boiling LN2 and stop pumping. 4) By the frequent interventions of the cryogenics operations crew. Occasionally the current of the magnet had to be reduced to levels which would have prevented the operation of the experiment. WEAR/MAINTENANCE The number of interventions and the cost for corrective maintenance increased with time and reached a peak in the second half of the 1980's. The type of maintenance intervention carried out during the experiment shut down periods between 1985 and 1995 are listed and plotted as illustrated in fig.1. We have defined 5 types of interventions which are: 1) standard (normal) maintenance, 2) replacement of piston labyrinth seal 3d stage, 3) replacement of piston labyrinth seal 1st, 2d, 3d stages, 4) remachining of 1st stage cylinders, 5) remachining of 1st and 2d stage cylinders. In fact not only the labyrinth seals of the piston underwent wear but also the cylinder walls. Therefore, in this period, almost every year the cylinder breech block had to be dismounted and delivered to Sulzer workshops for machining and repair. These were costly and cumbersome exercises. At each machining intervention the cylinder walls were reduced in thickness and its critical minimum value for the integrity of the mechanical structure was approached. As even these very frequent and costly maintenance operations were insufficient to prevent progressive loss of efficiency and guarantee the correct functioning of the plant in 1990, the replacement of the compressor by a better suited, modern machine, was envisaged. However, not only would this solution have been expensive (more than one million Swiss francs) it would have necessitated the interruption of the Omega's experimental programme for a full physics period while the new compressor was installed. Eventually, the problem was solved in a much cheaper and quicker way by inserting an auxiliary cooling stage between the compressor and the cold box. This to reduce the helium temperature at the suction side and at the delivery stage of the existing piston compressor.
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THE REFRIGERATOR, FUNCTIONING The R-22 Freon based refrigerator, AERMEC type RAX212, with cooling capacity of 50 kW, is driven by two Maneurop compressors of 9 kW electric input power each. For condensation of the R-22 Freon the refrigerator has a twin condenser battery (heat sink) with air heat exchangers of 132 m 2 cooling surface cooled by two air ventilators delivering 20000 m3/hr. The condensed Freon is expanded through thermally isolated evaporator/heat exchanger where the Glycol-water mixture is recooled. The refrigerator circuit is internally regulated by means of a thermostatic expansion valve which modulates the evaporation rate of Freon as a function of the thermal load. The water-glycol temperature is in turn controlled by an electronic, multistage, thermostat driven by a temperature sensor located at the input of the evaporator and acting on a three-way bypass control valve. The water-Glycol mixture is circulated with a pump. The hydraulic scheme of the system is presented in fig.2. The cooling system, i.e. the auxiliary refrigerator RAX and the two heat exchanger are installed in the compressor hall adjacent to the piston compressor at a distance of some 80 metres from the cold box. The auxiliary refrigerator RAX is in operation during the complete run of the OMEGA main helium cryoplant. The helium gas temperatures are regulated to the desired values as adjusted by the operators. PERFORMANCE Comparative measurements of the performance of the OMEGA refrigerator have been made, with and without the auxiliary refrigerator RAX, and observations of the benefits made. At nominal operation conditions, i.e. a fully charged magnet at 4800 A and at 25 ~ ambient, the temperature of the suction flow was cooled to 10 ~ while the high pressure delivery gas was at 13~ An increase in refrigeration spare capacity from 40 W to 120 W was measured, i.e. a net gain at the 4.5K level of 80 W. Also, due to the increase in mass flow, the re-injection rate to the third stage of the compressor could be reduced by 10 %. The gain on the 80 K level was approximately 350 W. At the same time the temperature of the LN2 circuit dropped from around 79 down to 76 K, thus reducing the risk of boiling nitrogen and failure of the pump. Just in case of necessity the RAX refrigerator could operate at much lower temperatures providing helium gas, for example, at close to 0~ This would further increase the performance of the OMEGA refrigerator. However, we have found that in normal cases this additional performance is not needed and the gas temperature is in general maintained between 8 and 12~ ECONOMICS/MAINTENANCE As can be seen in the graph of fig. 1, the number of interventions for maintenance was drastically reduced from the date of installation of the booster. In fact, in the years 1992, 93, 94 only standard maintenance was required. In the year 1995 a 3d stage labyrinth seal was exchanged in addition. The cost saving, due to greatly reduced maintenance, is noticeable, as is the ease of operation with reduced rates of troubleshooting which has also become an important factor. ACKNOWLEDGEMENTS We would like to thank D. Campi for his assistance in the selection of the auxiliary refrigerator. The important contributions of G. Cuccuru, M. Bongiovanni and F. Rodriguez to the success of this plant are acknowledged. REFERENCE Morpurgo, M., " The Design of the Superconducting Magnet for the 'Omega' Project", Particle Accelerators, Gordon and Breach, Science Publisher LTD, Glasgow, Scotland, 1970
Refrigeration System for the Atlas Experiment
F. Haug 1, j.p. Dauvergne 1, G. Passardi 1, D. Cragg 2, C. Cure 3, P. Pailler 3, C. Mayri 3, A. Yamamoto 4 1CERN - LHC Division, CH-1211, GENEVA 23, Switzerland 2RAL, Chilton, DIDCOT OX11 0QX, UK 3CEA, Saclay, Gif-sur-Yvette, F-91191, France 4KEK, Tsukuba 305, Japan
The ATLAS detector of the 27 km circumference LHC collider is of unprecedented size and complexity. The magnet configuration is based on an inner superconducting solenoid and large superconducting air-core toroids (barrel and two end-caps) each made of eight coils symmetrically arranged outside the calorimetry. The total cold mass approaches 600 tons and the stored energy is 1.7 GJ. The cryogenic infrastructure includes a 6
[email protected] K refrigerator, a precooling unit and distribution systems and permits flexible operation during cool-down, normal running and quench recovery. A dedicated LN2 refrigeration system is required for the three liquid argon calorimeters (84 m 3 of LAr). Magnets and calorimeters will be individually tested prior to their definitive installation at a large scale cryogenic test area. The experiment is scheduled to be operational in 2005.
INTRODUCTION ATLAS [1 ] will be one of four particle detectors designed for the exploitation of the LHC's capabilities for experiments with colliding protons and heavy ions. All of them envisage solutions with superconducting coils at 4.5 K [2]. The extent to which the ATLAS will use superconducting technology is unprecedented in complexity. This complexity is reflected in the associated helium cryogenic system. An equally complex refrigeration system is needed for the cooling of the three liquid argon calorimeters. The paper describes the cryogenic infrastructure to be installed at the LHC's collider point 1 at CERN, the refrigeration system and the operational scenarios we have studied and defined to date. The design of the ATLAS magnets and calorimeters is an intemational undertaking with contributions from laboratories in Europe, Japan and the USA. THE TOROIDAL MAGNET COMPLEX AND ITS CRYOGENICS The Barrel Toroid (BT) and the two End Cap Toroid magnets (ECT) produce a large volume toroidal magnetic field for the muon spectrometry. The BT is made of 8 rectangular coils with a length of 26 m and a height of 5 m. They are arranged around the central beam axis in the form of a cylinder with 19.5 m outer diameter [3]. The ECT's consist each of 8 rectangularly shaped coils housed in a common vacuum vessel of an outer diameter of 10.5 m. Peak field is 4 T at nominal current of 20 kA. The refrigerant flow is split in eight parallel circuits for each of the three magnet subsystems [4]. The cooling mass flow under normal operation conditions is ~600 g/s of two-phase helium for the BT magnets and ~300 g/s for each ECT circulated by means of a turbo pump backed up with a second identical one for redundancy [5]. For the BT a 5000 liter dewar and for each ECT a 1600 liter dewar provide sufficient autonomy for slow discharge in case of failure of the refrigerator. In case of fast discharge the stored energies of 1.1 GJ for the BT and 0.25 GJ for each ECT will be dumped in the cold mass of the magnets which heat up to 58 K (BT) and 53 K (ECT's) respectively. 111
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THE SOLENOID AND ITS CRYOGENICS The comparatively small solenoid of length 5.3 rn and inner diameter of 2.4 m is designed to provide an uniform magnetic field of 2 T at 8 kA for the inner tracker. Supercritical helium from common refrigerator is sub-cooled in a control dewar (250 liter) and expanded by using a J.T. valve to provide two-phase helium (7 g/s) close to full liquid phase. The liquid helium in the dewar also serves for secure slow discharge of the magnet. The solenoid is housed in the same vacuum vessel of the liquid argon barrel calorimeter to minimise the amount of material along the particles trajectories. THE HELIUM CRYOGENICS INFRASTRUCTURE A dedicated helium cryogenic plant and infrastructure is proposed to be installed at CERN's LHC point 1 (see fig. 1). It mainly consists of 1) at surface level: screw compressors, He storage tanks, a recuperation and purification system 2) at underground cryogenics service cavern: cold box, precooling unit 3) at main detector cavern: distribution system, local cryogenics for the four magnet subsystems. Based on the thermal budget of the magnets (see table 1) we presently foresee a refrigerator of 6
[email protected] and a compressor flow of 500 g/s (1-20 bar). Detailed studies are envisaged taking various scenarios into account such as precooling, baseline operation, recovery from fast discharge, permitting an optimisation of the thermodynamic cycle. The thermal budget of the four magnets is 2550 W of isothermal refrigeration at 4.5 K, 10.3 g/s of liquefaction for cooling of four pairs of current leads, and 10600 W for the thermal shields (feed 40 K, return 80 K). For the cooling of the magnets from 300 to 100 K, a dedicated precooling unit with a LN2/He heat exchanger will be installed. The 23 m3/d of LN2 required during cool down is supplied from two 50 m 3 storage tanks at the surface and the helium mass flow is withdrawn from the compressor/refrigerator circuit. COMPRESSORBUILOING SURFACE BUILDINGS
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Figure 1 3-D view of cryogenics areas of the ATLAS experiment; surface buildings, side cavern and main detector cavern. Cut out view of the detector showing the superconducting magnets and liquid argon calorimeters arrangement.
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Cryogenics parameters for the refrigeration systems of the ATLAS magnets and calorimeters. BT = Barrel Toroid Magnets, ECT = End Cap Toroid Magnet, ECC = End Cap Calorimeter.
Helium Cryogenics
Argon Calorimeter Cryogenics
Baseline operation Conditions BT 2 ECT Solenoid Liquid volume m3 Cold mass tons 350 214 5.5 Cold mass Shield tons 25 45 0.5 Stored energy GJ 1.1 0.5 0.04 Load 40 to 80K kW 6.3 4.02 0.28 Load4.5 K kW 1 . 2 4 1.22 0.09 Current leads g/s 3 6 1.3 Total equiv. 4.5 K kW 2.03 2.13 0.24 Refrigerator (with contingency)@4.5 equiv, kW Compressor mass flow (1-20 bar) g/s Cool down operation conditions
Total 15 570 71 1.7 10.6 2.55 10.3 4.50 6 500
Time 300/i00 K (pre-cooling unit) Time 100/4.5 K (refrigerator) (AT max = 40K) total Average cool-down power He mass flow pre-cooling unit (300-100 K) LN2 consumption
days
28
days
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days kW g/s m3/d
40 43 220 23
.
.
.
.
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Liquid volume Cold mass Isoth.load 89 K
Barrel 44 130
m3 tons kW
2 ECC 40 440
Total 84 570 19.1
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I 14
LN2 refrigerator
[kW
[25
Time 300 K/89 K (He/LN2 pre-cooler) (AT max = 40K)
total
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40
HELIUM DISTRIBUTION SYSTEM The refrigerants are distributed via a low loss transfer line system linking the precooler unit and the refrigerator of the side cavern to the cryogenics equipment of the main cavern. Our modular design of connecting the magnet subsystem as shown in fig.2 permits operational flexibility in cool down, baseline running and recovery modes. Feed and return flows of any defined temperature level at baseline load (supercritical He at 4.5 K, shield cooling 40/80K) form a closed circuit via a bypass at the end of the transfer line. For the precooling we foresee an independent line following the same principle. For reasons of thermodynamic optimisation, the enthalpy of the cold return gas is utilised in the refrigerator and/or the precooler (no heaters will be installed for warming up the helium gas flows). Pre-Cooler ,, .
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Figure 2 Helium Distribution System for the ATLAS Magnets. FUNCTIONING The normal cool down procedure foresees all magnet subsystem cooled in parallel from ambient to -~100 K by means of the LN2 precooler unit (limits on magnets: temperature gradients of 40 K and 2.5 K/h). From -~100 K to 4.5 K the refrigerator's J.T. flow will be used. However, sequential cooling of the magnets can be carried out. In both cases the total cool down time will be -~40 days.
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The modular design of the He distribution system permits the magnet subsystem to be run under different conditions if required. For example, one magnet could be in cool down mode from ambient while the remaining are already operating at 4.5 K. In another configuration one or more magnet(s) may be in quench recovery mode while the others are kept cold. Recovery time for any quenched magnet subsystem will not exceed four days. Various post-quench situations will be studied and optimal solutions investigated for different operational scenarios. THE LAR CALORIMETERS AND REFRIGERATION SYSTEM The three liquid argon calorimeters, with a total liquid inventory of 84 m 3 are: the barrel electromagnetic calorimeter (dimensions of vessel 4.5 m o.d., length 6.8 m) and the two end cap (both electromagnetic and hadronic) calorimeters 4.5 m o.d., length 3.3 m. The equipment of the dedicated LAr refrigeration system are located at 1) floor level: nitrogen compressor (for the LN2) refrigerator, LN2 storage tanks, helium compressor 2) the underground cryogenics service cavern: LN2 refrigerator (25 kW), precooling unit (He/LN2) 3) the main cavern: 2 LAr storage tanks each 50 m 3, LN2 buffer tank of 20 m 3, local auxiliary cryogenics. Precooling of the calorimeters from ambient temperature to 89 K in 40 days uses He/LN2 heat exchange. Helium at 1-3 bar is circulated with a compressor. At operational temperatures, the calorimeters are purged and filled with LAr delivered by truck from the surface area. Internal cooling of the LAr is done either directly with LN2 or indirectly with an intermediate LAr circuit (design decision pending) in horizontally heat exchanger tubes. The LN2 refrigerator provides the cooling for all operational modes (cool down, normal operation at 89 K) of the detectors. The 20 m 3 of LN2 in the main cavern is designed to give more than a day of autonomy in case of failure of the LN2 refrigerator. An additional back-up utilises the LN2 (2 x 50 m 3) from the tanks at the surface level. If necessary, the complete liquid inventory of any or all cryostats can be drained into the 100 m 3 LAr storage tanks near to the detector in the main cavern. THE ATLAS TEST FACILITY HALL All the cryogenic components must be tested prior to their final installation in the underground cavern. This will be carried out in a large experimental hall with 10.000 m 2 of surface area which will be transformed into a cryogenics test facility permitting individual tests of BT and ECT magnets and the three liquid argon calorimeters. Four test stands will be required for the BT magnets which will be operated in parallel. Helium precooling units both for the magnets and calorimeters will be provided. A helium cryoplant already existing at CERN with a capacity of 1200
[email protected] K will be used. The stringent schedule, especially the arrival of a pre-series prototype barrel magnet coil with approximately 1/3 of the length of the final magnets requires this test facility to be available in 1999, well before the start of delivery of the series magnets planned for 2001. REFERENCES 1 ATLAS collaboration, "ATLAS Technical Proposal", (~ERN/LHCC/94-43, Geneva, 1994 2 Bremer J., Dauvergne J.P., Delikaris D., Delmelle N., Haug F., Passardi G., Rieubland J.M., Kesseler G., "Cryogenics for CERN Experiments. Past, Present and Future", this conference 3 Baze J.M. et al., "Progress in the Design of a Superconducting Toroidal Magnet for the ATLAS Detector of LHC", 14th Int. Conference on Magnet Technology, Tampere, Finland, 1995 4 Cragg D., "The Cryogenic System of the ATLAS Experiment End Cap Toroids", this conference 5 Mayri C., Cur6 C., Duthil R., Cragg D., Haug F,.Passardi G., "Barrel Toroid Cryogenic System for the ATLAS Detector", this conference 6 Tanaka K., Yamamoto A., Doi Y., Haruyama T., Kondo T., Makida Y., "Cryogenic Design of the ATLAS Thin Superconducting Solenoid Magnet", this conference
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The Cryogenic System of the ATLAS Experiment End Cap Toroids
Derek Cragg Applied Science Department, Rutherford Appleton Laboratory, Chilton, Didcot, Oxon, OX11 0QX, UK.
The ATLAS Experiment proposed for the Large Hadron Collider will use a toroidal magnet system to achieve high efficiency muon momentum resolution. The End Cap Toroids (ECT's) are designed to provide bending powers in the range 4-8 Tesla-metres over the rapidity span 1.5-2.8 in the important forward/backward regions. Each ECT will have an outer diameter of approximately 11 metres, a length of 5 metres and a weight of 190 tonnes. They will each have eight separate coils and a single integral radiation shield which will all be contained in a common cryostat.
INTRODUCTION The muon spectrometer of the ATLAS general-purpose pp detector will be based on the configuration of large superconducting air-cored toroids shown in Figure 1 and will consist of a long barrel toroid (BT) and two inserted ECT's to generate the large field and strong bending power required. One ECT is shown withdrawn from its normal operating position to allow access to internal detectors. The Rutherford Appleton Laboratory will be responsible for the design of the ECT's [1] and this paper describes the cryogenic design.
END CAP TOROID WITHD RAWN
E N D CAP TOROID IN S E R T E D
BARREL TOROID
Figure 1 ATLAS superconducting toroid magnet system. The ECT's will be designed to provide bending powers in the range 4-8 Tesla metres over the rapidity span 1.5 - 2.8 in the forward/backward regions. While the BT design will be based on an open structure with 8 coils in individual cryostats the ECT design will be based on mounting 8 coils in a single large vacuum vessel approximately 11 metres in diameter and 5.6 metres in length.
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HELIUM REFRIGERATION SYSTEM The ATLAS helium refrigeration system will supply the BT, the two ECT's and a cemrally located solenoid. It will be designed to have the capacity to cool the whole ATLAS magnet system down in 40 days and to supply the peak demand required during base temperature operation, i.e. all the magnets charging at the same time. In order to give the necessary flexibility of operation, the system will be divided into two discrete components: a liquid nitrogen cooled pre-cooler and a helium refrigerator. The pre-cooler covers the temperature range from 300K to ~100K and will cool down and warm up the magnets: - both relatively infrequent processes. The helium refrigerator will cover the temperature range below ~100 K and will supply helium gas at 40K to the radiation shields and supercritical helium at---4.5K to the coil system during cool down and operation at base temperature.
CRYOGENIC DESIGN The main structural components of an ECT are shown in Figure 2. The current leads, which are not shown, will be installed vertically inside the services turret. SERVICES TURRET
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Figure 2 Main cryogenic components of an ECT. The heat load into the radiation shields will be minimised by using 30 layers of superinsulation. It will not be considered necessary to superinsulate the coil structure, only to reduce the emissivity of the surfaces by mechanically polishing the metal components and by covering the exposed plastic surfaces with aluminium foil. This has proved to be the most cost effective solution in the past for large cryogenic systems and will give a low value for the radiative heat leak into the system [2]. The current leads will be designed to take the high pressures developed during a fast discharge and will be cooled directly by using two-phase helium returning to the buffer/storage dewar from the coil cooling circuits. The cryogenic parameters of a single ECT cryogenic system are given in Table 1. Cool Down During cool down from 300K to ~100K it will be important to avoid thermally induced stress caused by imposing too high a temperature differential across the cold mass. The limits chosen for the ECT's are a cool down rate not to exceed 2.5 K/hour and a temperature differential not to exceed 40K. In order to simplify the control system and to avoid problems met by cooling the radiation shields and coils separately their cooling circuits will be connected in series during this period. The mass flow rate required from the pre-cooler to cool an ECT down from 300K-100K in 33 days will be --40 grams/second.
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Table 1 Cryogenic parameters of a single ECT. Parameter Cold mass Radiation heat load Conduction heat load from supports Coil charge/discharge Heat load from local cryogenics Maximum heat load from helium pump Operating current Current lead helium consumption Refrigerator demand at base temperature Helium pump mass flow rate
Radiation Shield 22500 kg (22.5 tonnes) 1380 W 330 W 75W
Coil 107000 kg (107 tonnes) 300 W 29 W 56W 25W
-
125 W
0.0086 kg/s (8.6 gm/sec) -
20000 A 0.003 kg/s (90 litres/hr) 0.028 kg/s (28 gm/sec) 0.28 kg/s (280 gm/sec)
Below 100K these thermal restrictions will not apply so it will be possible to connect the radiation shields and the coils directly to the refrigerator, cooling down the shields with 40K helium gas and the coils with supercritical helium. The coils will be cooled in the same way during the recovery from a fast discharge. If the helium refrigerator capacity needed to run the ECT coils at base temperature is used to cool them down from ~100K and recover them from a fast discharge the cool down times will be ~7 days and ~2 days respectively. Base Temperature Operation The radiation shields will be designed to run with a temperature differential across them of 40K, i.e. they will return helium to the refrigerator at a temperature of 80K. In order to avoid a fast discharge when a refrigerator fault occurs, the coils will be cooled during base temperature operation by pumping two-phase helium round the cooling circuits from a buffer/storage dewar. The two phase helium will be produced by expanding supercritical helium from the refrigerator through a JT valve. Flow instabilities in the coil cooling circuits will be prevented by pumping the twophase helium at a mass flow rate which will return it to the dewar with a liquid content of not less than 90% (quality factor 0.1).
CRYOGENIC SYSTEM The proposed layout of the cryogenic system is shown in Figure 3. Internal Cryogenic System The internal cryogenic system of an ECT will not contain a significant volume of helium and will consist simply of the current leads and the cryogenic cooling circuits of the radiation shields and the coils. Local Cry.ogenic System The local cryogenic system of an ECT will consist of a valve box, a buffer/storage dewar, two helium pumps and the connecting transfer lines. Because space within the experimental cavern will be at a premium and high magnetic fields will be present in the detector region, the valve boxes, buffer/storage dewars and the helium pumps will be placed at the wall of the cavern on a high level gantry. The proposed flow diagram of the local cryogenic system for a single ECT is shown in Figure 4. The valve box will contain all the control valves required to cool down/warm up and run an ECT at base temperature and all the pressure relief valves needed to protect the system during fault conditions and fast discharges of the coil. In the event of a refrigerator failure the buffer/storage dewar will have sufficient capacity to enable the coil to be discharged from full current in about 2 hours.
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TOROIDS
REFRIGERATION PLANT
BUFFER DEWAR AND HELIUM PUMPS
END CAP TOROIDS
TRANSFER LINES
GROUND LEVEL
VALVE BOXES
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Figure 4 Flow diagram of the local cryogenic system for a single ECT. The two helium pumps will be proprietary centrifugal pumps and each will have the capacity necessary for normal operation. One will be on line while the other will be held in reserve should the first develop a fault. Since the ECT's will need to be moved out of their operating positions during shut down periods they will be connected to their respective valve boxes by flexible cryogenic transfer lines. All the other transfer lines will be of the conventional rigid, tube-in-tube variety.
REFERENCES The AECT reference design report, RAL report AECT/5 5/96 (1996). Cragg, D., Thermal and vacuum insulation for large cryostats, RAL report AECT/56/96 (1996).
Cryogenic Design of the ATLAS Thin Superconducting Solenoid Magnet Ken-ichi Tanaka, Akira Yamamoto, Yoshikuni Doi, Yasuhiro Makida, Tomiyoshi Haruyama, Takahiko Kondo, National Laboratory for High Energy Physics, KEK 1-10ho, Tsukuba, Ibaraki, 305, Japan Cryogenic characteristics of a thin superconducting solenoid magnet has been studied for a high energy particle-detector, ATLAS, in the Large Hadron Collider (LHC) project at CERN. The thin solenoid wound with aluminum stabilized superconductor is indirectly cooled by forced flow of two-phase helium in serpentine cooling tube on the outer support cylinder of the coil. This report describe a cryogenic design of the thin solenoid magnet and its cooling system. INTRODUCTION A thin superconducting solenoid magnet is planned to be developed for the ATLAS detector which is one of major particle detector systems in the Large Hadron Collider (LHC) project at CERN [1, 2]. It is designed to provide a central magnetic field of 2T in a warm-bore volume of 2.3 m in diameter and 5.6 m in length for precise momentum measurement of secondary particles produced in 14 GeV proton-proton head-on collisions in the LHC accelerator. Since the solenoid is required to be as thin ( and transparent) as possible in terms of radiation length for particles to traverse the solenoid magnet wall with minimum interaction. The solenoid coil is wound with aluminum stabilized superconductor on inner surface of an outer support cylinder, and is cooled indirectly by using two-phase helium flow in serpentine cooling pipe welded on outer surface of a coil support cylinder. The resource cryogen is supplied by a common large refrigerator for whole ATLAS detector magnet system consisting of toroidal magnets and a solenoid magnet [3,4,5].This report describes a cryogenic design of the ATLAS solenoid and associated interface to the expected common refrigerator system for the ATLAS superconducting magnet system. SOLENOID COIL AND CRYOSTAT The solenoid coil consists of a single layer-coil wound with aluminum-stabilized superconductor and an outer support cylinder as shown in Fig. 1. The coil is directly wound inside the support cylinder made of high-strength aluminum alloy. A total cold-mass thickness of 43 mm resulting in a cold mass of 5 tons has been determined with an optimization of E/M (stored energy / cold mass) ratio of 8.4 kJ/kg [6]. It enables to absorb a full stored energy of 42 MJ into the cold mass with an averaged coil temperature rise of about 80 K, in case of quench. The temperature rise may be homogenized with fast thermal propagation given by axial pure-aluminum strips placed on inner coil surface [6]. Major coil design parameters are summarized in Table 1. Fig. 2 shows an isometric view of the coil structure. A cooling pipe with an inner diameter of 18 mm is welded to the outer surface of the outer support cylinder. The single serpentine path has 24 axial passes with a circumferencial pitch of 33 cm. The cold mass is supported by 12 triangle-supports made of glass fiber reinforced plastic (GFRP) at each axial end of the coil structure. The triangle supports are designed to provide combined functions in supporting against the expected maximum load within 3 G. A coil end is axially fixed by the support and the other end is axially free to allow axial thermal shrinkage according to the coil cool-down. The radial shrinkage may be allowed by radial tilting of all triangle supports. .... Table 1. Design parameters of the ATLAS solenoid. Coil Inner Radius 1.218 m Coil Half Length 2.65 m Central Field 2.0 T Peak Field 2.6 T Nominal Current 8,000 A Stored Energy 42 MJ FJM Ratio 8.4 kJ/kg 119
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Fig. 2. An isometric view of the cooling tube and the coil support scheme. To save material to be used in the cryostat, the coil is installed into a common vacuum vessel with another cold detector component, Liquid-Argon Calorimeter (LAr-Cal) and its cryostat. No outer radiation shield is provided because of the inner wall of the Liquid-Argon calorimeter cryostat which may provide a thermal boundary at 80 K. The inner radiation shield is provided for an intercept against thermal radiation from the inner vacuum warm bore tube. As shown in Fig. 3, the cryogen lines and the superconducting leads of the solenoid extend through other detectors in an 10 m long chimney to a control dewar, which is located outside top of the muon detector at the axial detector-center. The control dewar provides optimized cryogen flow into the solenoid magnet in various cooling modes, and supplies cold gas flow to current leads passing through the control dewar. CRYOGENIC DESIGN Thermal loads In a steady state, the net thermal load into the ATLAS solenoid magnet is estimated to be about 50 W at 4.2 K and about 450 W at a radiation shield temperature of 6 0 - 80 K. An eddy current loss of about 20 W in the support cylinder is additionally required during a magnet charging/discharging period of 20 rain., and a cold helium gas flow of 1 g/s is required for a set pair of current leads during the solenoid excitation. For a purpose of liquid helium level control, a heater of about 10 W may be consumed in the control dewar. As a result, a total thermal load of 100 W including a contingency of 20W is considered in the coil and the control dewar at 4.2 K. A summary of the thermal load is given in Table 2. Table 2. Thermal load of the ATLAS thin solenoid. 4.2 K 80 K 50 W (coil) 10 W Radiation 50 (chim. & C/D) 5 50 (coil) 5 Conduction 250 (chim. & C/D) 30 20 Eddy current loss (coil) (C/D) 10 Control heater Current leads (C/D) 20 100 Contingency 100 W 500 W
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Cooling scheme and parameters Figure 4 shows a planned schematic flow diagram for the ATLAS solenoid cooling system. Table 3. gives major cooling characteristics for the solenoid. The helium gas as cryogen resource is given by a common large refrigerator and a pre-cooler system for the ATLAS magnet system at CERN [4]. The solenoid is precooled by using a cold helium gas flow of about 20 g/s, supplied by the pre-cooler under safety constraints given in Table 3 to eliminate excessive thermal stress during a transient pre-cooling period. Depending on the operational condition, warm helium gas may be mixed with the cold gas, as an optional mode, to ensure the temperature control of the cold gas. In a steady state operation at 4.4 K, the two phase helium flow is supplied to keep the coil temperature at about 4.5 K( 35g/s is required. Using injector 2 can increase mass flow up to 40g/s. If injector 1 is on line, pressure in the subcooler may decrease to 0.05MPa, the temperature of outlet helium flow falls to 3.6K, but the mass flow through TM has to be reduced. According to our experience large mass flow is more effective for the HT-7 TM stability than supplying lower temperature helium flow. The 80K copper shields in the cryostat are cooled with LN2 fed from a lm3 storage vessel. The height of the vessel is 2m above the tokamak. For plasma chamber baking regime the LN2 is driven by a pump. According to the LN2 consumption rate the shield heat load is about 3.5kW/80K. The two LN2 plants can produce 300L/hr each. COOLDOWN The HT-7 TM (--14000kg weight) has been cooled down five times successfully for the primary test, engineering tests and normal plasma experiments. It takes a total of 90-100 hours from room temperature to 4.6K. The cooldown speed is limited by three factors: the helium flow temperature difference between TM outlet and inlet < 70K to avoid mechanical stress damage in TM, the supply helium pressure < 1MPa to ensure the electrical insulator safety on the helium supply tube lines in the cryostat, and the refrigeration capacity. When the return flow temperature is above 200K the mass flow
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is less than 17g/s and is driven by a small compressor. While the temperature below 200K a big compressor instead of the small one. A LN2 pre-cooler in the refrigerator provides refrigeration until return gas temperature falls down to 120K. When the return flow temperature reaches 120K four expanders start and cool the heat exchangers in the refrigerator. While the mass flow temperature falls to < 100K the expanders provide refrigeration and make supply gas temperature < 80K. From 80K to 4.6K it takes 7 hours only. When the return flow temperature reaches 6K the flow enters a 250L LHe vessel in the refrigerator. Then LHe from the tokamak will accumulate more and more in the vessel within a few hours thus the TM is cooled down completely and ready for charging. The thermal radiation shields (--8000kg) in cryostat is pre-cooled with cold nitrogen gas until the maximum temperature of the shields falls to 120K, then LN2 is supplied to continue cooling the shields. The shield cooldown speed must be similar with TM and be carefully controlled. Fig.3 shows a typical cooldown process of TM and shields.
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Fig.2 Flow diagram HEAT LOADS AND STEADY STATE OPERATION When TM is cooled down to 4.6K the vacuum in cryostat should reached --4xl0-4pa if the leak level is normal. The total heat load of TM is about 130W/4.6K. To cool a couple of 7000A current leads 0.7g/s mass flow from the return line is required. The heat loss on the transfer line between the refrigerator and current lead block is estimated total 20W including 12m-length transfer tube, five bayonet connections and a bypass valve. The detail items of heat load are listed in Table 1. Because the 48 NbTi/Cu superconducting strands embedded in HT-7 conductors are parallel arrangement without twist. When magnetic field changing the coupled eddy currents can cause TM premature transition to the normal state. So to ensure TM steady operation a mass flow of 35-40g/s is required and the current ramping speed is limited while the toroidal field >0.5T. Running two expanders and a big compressor can back up the tokamak steady operation at the toroidal field < 1.6T. Injector 2 on line is necessary to increase mass flow. When the toroidal field > 1.6T a small compressor must run to enhance refrigeration and mass flow. At the mass flow of 35-40g/s the inlet pressures are 0.24-0.26MPa, and the return pressure is 0.13-0.135MPa. So the upstream flow is
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supercritical and the downstream is two-phase flow. Fig.4 illustrated a T-S diagram for the steady state operation mode of refrigerator No.1. The parameters of pressure, temperature and mass flow are marked in the diagram. Table 1 Heat loads Thermal Radiation Gas Conduction at 5xl0-apa Conduction through Support Structure and Measurement Wires 50 Soldered Joints between the Windings at 6000A Current Mass Flow Cooling the Current Leads Transfer Line
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turbo-expander T3 was installed just before J-T valve to increase the cooling capacity from 4 kW to 8 kW. The helium compressor unit consists of 6 oil flooded screw compressors and 4 stage oil filters. Liquid helium produced in a 12,000 L liquid helium dewar is distributed to 16 SCC cryostats installed in about 200 m long straight section of underground TRISTAN tunnel through large size about 250 m long multi-channel transfer line. The liquid helium is also supplied to SCC test stand where we could cool down and test the spare SCC in 2 cryostats. Each cryostat contains two 5-cell 508 MHz SCC made of about 2mm thick pure niobium sheet. The cold mass and amount of liquid helium stored in a cryostat are about 1,000 kg and about 830 L respectively. The total amount of liquid helium handled in the system is about 16,500 L. For the helium gas recovery 9 x 100 m 3 medium pressure (1.9 MPa ) helium gas storage tanks are connected to the system. A 6.5 kW at 80 K liquid nitrogen circulation system with a turbo-expander and 2 screw compressors supply liquid nitrogen to 80 K thermal shield in the 16 cryostats in parallel through mutichannel transfer line. A off-line helium gas recovery system consists of a 5 stage air-cooled oil lubricated reciprocating compressor (15 MPa, 150 Nm3/hr ), high pressure low temperature purifiers (80 K, 15 MPa, 150 Nm3/hr ), and high pressure storage vessels (4 x 1350 Nm3). The whole cryogenic system is controlled by means of a process control computer system. MAINTENANCE The inspection and maintenance of the cryogenic system were performed regularly once a year during summer shut down of the system. The cryogenic system is obligated to inspect the gas tightness of the system, to measure the thickness of the vessel and pipe walls, to check the operation of safety valves and to correct pressure gauges and thermometers once a year by high pressure regulation law. During this period mechanical machines with wearing parts, such as the compressors, including the air compressors, motors, oil pumps, turbines and valves were inspected carefully and repaired if necessary. The charcoal adsorber of the main helium compressor unit was also changed. In addition the maintenance of the process control computer system were carried out. Compressors Figure 3 shows the history of the replacement and the lifetime of the mechanical shaft seals of the compressors. At an early stage in the operation the mechanical seals of the compressors had to be replaced due to leakage of oil from it. These replacements are shown by black bars in Fig. 3. The oil leaka__~a~!e ...... was
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caused by worn out of pins used for fixed the seal by the mechanical vibratioo of me compressors. These failures were overcome by improving the hardness of the pins. After the adoptio,, of new type pins lifetime of the seals increased as shown in Fig. 3. The mechanical shaft seals of the main compressors were replaced at annual regular maintenance for precaution. These replacements are shown by white bars. The compressors of the 2nd stage were send to the factory for detailed inspection including the beating every three years. The checks revealed no damage. Oil Adsorbers The oil adosorber consists of 4 stages oil removal filter : mesh demister coarse oil removal just behind the compressors, 2 fine oil coalescence filters (BALSTON BX) and charcoal adsorbers. Typical measured values of the oil concentrations of outlet gas from the 1st, 2nd and 3rd were 4,000 ppm, 63 ppm and 2.5 ppm in weight. The oil recovered at the oil filters are recycled into the systems. The oil adsorber 4th filters remove the rest of oil aerosols from the helium. The adosorber of the main compressor unit is filled with 1580 L of charcoal and 400 L of molecular sieve. The oil concentration of the outlet gas from the adosorber was less than 52 ppb in weight. The charcoal adosober had to be changed every year (about 7,000 hours Table 1. List of fatal failures during 7 years operation of TRISTAN SCC cryogenic system No.
.
.
.
4. 5. 6. 7. 8. 9.
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Item
Da~
Failure (Action)
Cold Box Low bearing gas pressure of T2 turbine Nov.7, 1988 (Restart) Helium Compressor (C3) Oil leakage at the mechanical seal of the compressor unit Dec. 29, 1988 (Replace of mechanical seals) Helium Compressor (C2) Oil leakage at the mechanical seal of the compressor unit (Replace of mechanical seals and Restart) Jan. 14, 1989 High discharge pressure of the compressor unit Helium Compressor (Restart) Jan. 26, 1989 High temperature of oil due to failure of the cooling water Cooling Water [Repair of cooling water unit and Restart] Jan. 28, 1989 Electric power outage due to thunderstorm Electric Power (Restart) May 3, 1989 High temperature of oil due to incorrect set of thermal Helium Compressor switch for interlock Jun.22, 1990 (Set correct value and Restart) Erroneous removal of a control relay during the operation Cold Box (Restart) May 29, 1991 High discharge gas temperature due to incorrect set value Helium Compressor of thermal switch for interlock Jul. 8, 1991 (Set correct value and Restart) Turbine T1 trip due to excess amplitude of radial vibration Cold Box (Restart) Oct. 1, 1991 Electric power outage due to failure at transformer room Electric Power (Restart) Mar. 25,1992 Low pressure of instrument air due to valve failure Air Compressor (Repair and Restart) Feb. 11, 1993 Malfunction of unloaders due to failure of controller Helium Compressor (Manual control of unloaders and Restart) Nov. 25, 1993 Electric power outage due to failure at substation Electric Power (Restart) Mar. 24, 1994
Downtime Cryo. System Beam 4 hr + 36 hr*
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Fig. 4 Statistics of the failures and fatal failures during 7 years operation of the system
intervals), because the guaranteed lifetime of the charcoal adosorber is 8,000 hours. The replaced new charcoal filter was activated by removing the water using 120 ~ dry nitrogen gas. FAILURES OF THE CRYOGENIC SYSTEM The total number of failures in the cryogenic system during about 7 years operation was 85. This number includes 71 failures which did not interrupt the operation of the system and 14 fatal failures which caused the whole system stop. Figure 4 shows the statistics of failures and fatal failures as the function of the operation hours. We classified the failures into seven categories : 1) helium refrigerator (coldbox, turboexpanders, distribution system), 2) compressors (mechanical seals, oil pump, oil valves etc.), 3) computer control system, 4) liquid nitrogen circulation system, 5) cooling water, 6) instrument air, 7) electric power (main power failure by thunderstorm etc.). The fatal failures and down time of the cryogenic system and the beam of TRISTAN are summarized in table 1. In the table * indicate the recovery time i.e. time require the cryogenic system recovered to the same condition as just before the system down. SUMMARY The cryogenic system for the TRISATN superconducting RF cavity was operated very stably and reliably for about 38,000 hours in 7 years. The experience in long operation of the large cryogenic system shows the compressor is the key component and the regular maintenance of the system including utilities is essential to attain the system reliability. ACKNOWLEDGMENT The authors wish thanks Professors S. Kurokawa and Y. Kimura for thier continuous support and encouragement and the operation crew, the staff of the superconducting cavity group and Mr. M. Noguchi of Mayekawa for their devoted support and many helpful discussions. REFERENCES 1. Y. Kimura, TRISTAN project and KEK activities, in: Proc.XIII the International conference on high energy accelerators", Novosibirsk, U.S.S.R., (1986) 2. K. Hara et al., Cryogenic system for TRISTAN superconducting RF cavity, in: "Advances in Cryogenic Engineering", Vol.33, Plenum Press, New York (1988),p.615 3. K. Hosoyama et al., Cryogenic system for the TRISTAN superconducting RF cavities: performance test and present status, in: "Advances in Cryogenic Engineering", Vol.35, Plenum Press, New York, (1990) p.933 4. K. Hosoyama et al., Cryogenic system for TRISTAN superconducting RF cavities: upgrading and present status, to be published in: "Advances in Cryogenic Engineering", Vol.37, Plenum Press, New York,(1992)
Large scale refrigeration
Compressors
This Page Intentionally Left Blank
High Power Refrigeration at Temperatures Around 2.0 K
Guy Gistau-Baguer Air Liquide, Advanced Technology Division, BP 15, 38360 Sassenage, France
The possible technologies for pumping on the liquid helium bath are reviewed. An emphasis is put on cryogenic dynamic compressors, their bearings and their wheels. Arrangement of the compressors and their behaviour during transient situations are analysed. Information on existing machines is given. I thank companies which provided me with data concerning their machines. However, I may not be aware of all developements or applications and some wrong information may have crept into my paper.
INTRODUCTION There are very good reasons for operating magnets and resonant cavities at temperatures lower than T(k) (2.17 K). The saturation curve of helium shows that the pressure on the bath is 3129 Pa (0.03129 b) at 2.0 K and 1638 Pa (0.01638 b) at 1.8 K. This means that the theoretical compression ratio to reach atmospheric pressure is between 32:1 to 61"1. When taking real pressure drops into consideration, the tree compression ratio is higher, and can reach 100:1. Plants of power higher than 200 W are considered. HOW TO GET TEMPERATURES AROUND 2.0 K ? In order to achieve the required high compression ratios, pumping on the liquid helium bath can be made : 1) at room temperature, 2)at cryogenic temperatures or 3)partially at cryogenic then at room temperatures (see figure 1). Possible technologies are of the volumetric or dynamic type : see Table 1. At room temperature, all machines which are used for helium compression have to cope with its subatmospherlc \ ~ ~oots
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Table 1. Possible technologies for operation at temperatures around 2.0 K (legend : best suited, low compression ratio, limited volumetric flow rate) ~_-==_
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Ejector Radial Axial-radial
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Screw significant heat of compression. Rotary vane pumps have limited volumetric flow and they must be specially cooled. The Roots pumps are well suited for high compression ratio when processing air at very low mass flow rate, but helium operation actually limits the pressure difference to about 3000 Pa (0.030 b). Roots pumps have been used in a number of situations, either alone, with up to 5 stages in series by BOC (UK) [ 1], or in combination with rotary vane pumps : 3 Roots in series followed by 1 stage of 3 rotary vane pumps in parallel by Leybold (Germany) [2]. Existing piston and scroll compressors have insufficient volumetric flow. The ejector has insufficient compression ratio [3]. Liquid cooled machines like liquid ring pumps or oil lubricated screw compressors can withstand reasonable compression ratios. The reasonable suction pressure of liquid ring pumps is about 3000 Pa (0.030 b), therefore they must be used in combination with a first stage which can be a Roots. They are operated with the same kind of oil that is used in the screw compressors. Up to now, screw compressors have only been used for operation at higher temperatures [4]. They also need a booster stage. Room temperature pumping has the great disadvantage of having part of the circuits 9piping and machines, operating at sub-atmospheric pressure in the air. It is very likely that air leaks will appear and plug cold parts, which is not compatible with long term operation. It is therefore necessary to incorporate purification means which can be regenerated during operation of the plant, such as a dryer at the inlet of the cold box and adsorbers inside it. The heat exchangers operating at very low pressure are difficult to build : they must have a low pressure drop in order to keep the size of the first stage pumping machine reasonable, but have also to be efficient. Helium compression can also be made partially at cryogenic temperatures and partially at room temperature [5]. The heat exchanger becomes easier to design and make because its operating pressure is higher. For operation at cryogenic temperatures, a piston compressor has been developed by CCI and Fermilab (USA) [6] aiming at a temperature around 3.5 K, but volumetric flow and reliability are limited. Today, no system of this type is in operation. The cryogenic ejector, which is a tricky device, has insufficient compression ratio [7]. A scroll volumetric cryogenic compressor is presently been developed by the French Atomic Energy Commission [8]. The concept seems promising in that it shows its reliability. Finally, for temperatures around 2.0 K, cryogenic dynamic compressors are the most commonly used machines at the present time, which warrants a more-detailed analysis. Dynamic compression" a reminder Figure 2 is a customary method of presentation of a typical performance map for a compressor [9,10] : the pressure ratio across the whole machine is plotted as a function of (( reduced ~ flow 9rh * T ~ / P (rh 9mass flow rate, T 9inlet temperature, P 9inlet pressure), for fixed values of ~( reduced ~ speed " N / T ~ (N : rotational speed). At the left upper limit of the map, the stall line shows the limit of stable operation, unstable operation being characterised by a severe oscillation of the mass flow rate through the machine. The fight part shows the choked region where the constant speed lines tend to the vertical. No further increase in reduced flow is possible since sonic velocity is attained through some section of the machine. The efficiency hill can also be plotted on the compressor map. The characteristic number" k, = co * Q0.5 / AHO.75(co 9rotational speed, Q~ 9inlet volume flow, AH" enthalpy difference), is used as an index to flow path shape and type of machine in a range of applications. A low number means that the energy change is high, relative to the flow rate (high compression ratio) 9a radial impeller is needed. A high number means that the energy change is low, relative to the flow rate (low compression ratio) : an
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Figure 3. Cross section of a cryogenic centrifugal compressor with active magnetic bearings
Figure 2. Overall characteristic map of a dynamic compressor
axial impeller is needed. The radial wheel structure is best suited to high compression ratio which is what we are seeking. However, in the case of a cryogenic compressor, there are other parameters to be taken in consideration. For example, an axial-radial wheel has a lower diameter than a radial one ; therefore, phenomena like heat inleak and energy dissipation due to friction, which are both linked to the overall dimensions are reduced, sometimes giving a higher isentropic efficiency. The bearing system The bearing system required for a dynamic compressor must have similar characteristics to an expansion turbine : high rotational speed, no pollution, high reliability, low maintenance. Present machines incorporate 91) ball, 2) gas and 3) magnetic bearing types. The improvements to ceramic ball bearings has given running speeds compatible with operation of dynamic compressors. The life time is somewhat limited, but current expectations are in the order of a few thousands operating hours between maintenance. Linde/PBS (Germany/Czech Republic) produced such a machine operating at 1000 Pa (0.010 b) suction pressure [11]. Barber Nichols (USA) also produces such machines which operate at higher suction pressure [12]. Dynamic gas bearings are used in machines which are operating at suction pressures not lower than 0.05 MPa (0.5 b), therefore, the operating pressure of the bearing is about atmospheric. ]HI (Japan) produced such machines with foil bearings for FNAL [13]. However, there is a concern about the load bearing capability of the bearing system, when the suction pressure of the compressor is as low as 1000 Pa (0.010 b). Static gas bearings and a turbine drive are used in a cryogenic compressor developed by Air Liquid: for CERN. It seems to be a promising concept, and it is not yet in the final phase. The problem of the leaktightness between the bearings and the cycle has been solved by means of a labyrinth, a special chamber of which is connected to a mechanical pump. Magnetic bearings can operate at room or cryogenic temperature (see figure 3). They incorporate an electrical drive. Up to now, the active magnetic beating cryogenic compressors are the only type to have been "industrially" used in large scale refrigeration at 2.0 K or lower (see Table 2). Table 2. Comparison of bearing systems
Oil Ball Gas static Gas dynamic M:.::_:+a~netic
speed speed no low pressure
::::::
labyrinth labyrinth
? good good H i
g
h
High 8000 h ? High High ....................
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CRYOGENIC COMPRESSOR ARRANGEMENTS. The simplest cryogenic compressor arrangement is the pure series one, as used for CEBAF. The last stage discharge temperature is about 30 K for a 2.0 K operation and as high as 40 K for a 1.75 K operation. Operation of the last stage requires a high rotational speed because of the high inlet temperature and the small volumetric flow rate. The "high" discharge temperature makes it essential to connect the discharge side of the last stage compressor to the correct temperature level in the main cold box heat exchangers. If the design of the last stage is difficult, due to excessive power or rotational speed, or when a lower discharge temperature is sought, interstage cooling becomes necessary. It can be achieved by means of some boiling fluid or gas withdrawn from the cold box. Decreasing the suction temperature of a cryogenic compressor reduces the size of the machine and its power input, but such cooling power is still provided by the refrigerator. Therefore, the optimum efficiency of the system requires a delicate trade-off. Interstage cooling must be compatible with transient situations like cooldown. Different arrangements are possible. Up to now, none of them have yet been incorporated into a refrigerator. THE COOLDOWN PROCESS The cooldown process consists of changing the overall compression ratio from 1"1 up to the nominal value i. e. 40:1 or 100:1. Volumetric machines are slightly affected by a change in compression ratio within their operating range : when it increases, the volumetric efficiency is reduced. Therefore, the cooldown process is an easy operation. In the case of dynamic machines, it is necessary to keep the operating point of each of the stages inside the stable region of the compressor map, keeping in mind the peculiarities of a cryogenic compressor" the operating temperatures and pressures at the beginning of the cooldown process are different from the nominal operation ones, the mass flow is the same for each machine (except if partial bypasses are used). Consequently, the only parameter which is free is the rotational speed of each stage [ 14]. If we compare with the usual problem of starting a train of compressors operating at room temperature, we can identify the differences : the temperatures during the startup process do not differ very much from the nominal ones, therefore recycling is easier, but generally, the rotational speed of all stages is the same. If at least one stage of volumetric machine is incorporated downstream of the dynamic compressors, it makes the situation easier because the almost constant volumetric flow of the volumetric machine allows the dynamic ones to operate with a compression ratio which is reducing with flow, so, the operating point of the dynamic compressor stays in the correct region of the map. TRANSIENT SITUATIONS The equipment to be cooled at temperatures around 2.0 K generate heat loads which are lower during periods when the system is not operating at nominal. During turn-down situations it is generally required that the pressure on the helium bath is kept constant, so the compression ratio is to be kept constant when the mass flow rate is lower. For volumetric machines, there is no problem. However, the situation is delicate with dynamic compressors 9the operating point on the map is moving to the lett and hits the stall line. One easy, but thermodynamically non satisfactory, way is to compensate the missing flow by injection of electrical power into the liquid bath. H. Quack proposes to allow the discharge pressure of the dynamic compressors to decrease with the mass flow [ 15], in combination with a temperature Table 3. Comparison of pumping methods ..................................................................................................
Cryogenic power Sub-atmospheric circuit Low pressure heat exchanger Stability Reduced flow operation ..Ma!ntenanc_e
..........................
limited yes delicate natural natural
................................................................ .!mpo
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unlimited yes easy easy easy
ant ................................................. r e d u c e d
o.
..e..nic .....
unlimited no none reasonable delicate
............................................ l o w
...................
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stabilisation at the inlet of the third stage compressor. The disadvantage is to have part of the room temperature system operating at sub-atmospheric pressure. CERN proposes to operate at a lower first stage suction pressure [ 16]. There is still a significant amount of work to be done on this topic. EXISTING MACHINES, PROJECTS Today, few large refrigerators are operating. The Stanford refrigerator was the very first one. Let's imagine what a challenge it was at that time ! Then, 3 similar machines were built by Messer Table 4. Existing machines and projects. (legend ; R" Roots, CC 9cryogenic compressor, V 9vane pump, LR" liquid ring pump, s 9series, p 9parallel) Temperature Power
Pumping system
.............................................................................................................................................................
CERN BNL CERN TTF Tore Supra CEBAF
1.8 K 2.6K 1.8 K 1.8 K 1.75 K 2.0 K
300 W 13700W 360 W 200 W 320 W 4800 W
1.8 K 2.0 K
2400 W 3720 W
Efficiency vs..Camot
5 Rs 4CCs 1 CC + 3 Rs + 3 Vp 2 Rp + 2 Rs + 4 Vp 0.140 2 CCs + 2 LRs 4 CCs 0.185
Elapsed
....... a t . .......
time Remark ...........................................
not operating not operating
22700 h 16000 h
avail. > 98%
Projects
LHC TESLA
x 8 plants x 14.pl.a..nt.s........
Griesheim, Linde and BOC. All of them were based on room temperature pumping with Roots. Today, none of them are being operated at the original design temperature. The large plant at Brookhaven National Laboratory (USA) : 13.7 kW @ 2.6 K, had a very innovative design, incorporating different cryogenic compressors and circulators with room temperature oil bearings [ 17]. Unfortunately, the accelerator which was to be cooled by this refrigerator was cancelled and consequently, as the refrigerator had some difficulty in reaching its nominal performance, there was no further attempt to fix the situation. The first plant to have been "industrially" operated is Tore Supra : 320 W @ 1.75 K incorporating cryogenic compressors (with cryogenic active magnetic beatings) and liquid ring pumps. Operational experience is good : since 1986, 22700 hours of operation at temperatures lower than T()~) have been accumulated [18]. High reliability of cryogenic compressors and liquid ring pumps has been demonstrated. The second large refrigerator was built for CEBAF : 4800 W @ 2.0 K, where the pumping system is totally cryogenic. After a quite long startup period, experience has been really positive : since the start of the operation of the accelerator 16000 hours have been accumulated at temperature around 2.0 K during runs of more than 80 consecutive days and the total availability of the cryogenic system is more than 98 % [19]. The CERN LHC cryogenic test stand is a room temperature pumping system (Roots and rotary vane pumps) for 120 W @ 1.8 K which can be boosted with a cryogenic compressor in order to triple the power. CERN is presently evaluating two cryogenic compressor technologies 9ball and static gas beating [8] in view of the LHC project. The Tesla Test Facility (TTF) at DESY Hamburg (Germany) 9 200 W @ 1.8 K, is a room temperature pumping system, very smilar to the CERN one. There is also a project for boosting the present system with a cryogenic compressor. Current active projects are LHC at CERN and TESLA. Both require a large amount of power 9about 20 kW @ 1.8 K for LHC and about 52 kW @ 2.0 K for TESLA. For both projects, extensive studies by laboratories and industry are being conducted in order to decide what would be the best suitable structure for the refrigeration system. In case of LHC, one has to take into consideration that the existing large LEP refrigerators will be part of the LHC system. It seems that a prudent approach could be to build and test full-size sensitive components before incorporating them into the system. Such full-size tests would allow to experience the operation of the whole refrigeration system, i.e. the pumping system and the refrigerator. All such tests should be conducted with test equipment which can be included into the cold box to simulate the steady state and transient situations.
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CONCLUSION. Today, refrigerators operating at temperatures around 2.0 K are demonstrating their capability to cool large cryogenic systems in an way. This is a very important and positive input for projects as LHC and TESLA the size of which is about one order of magnitude larger. However, there is still some thought to be given to the problem of operating such refrigerators at reduced capacity in an optimized thermodynamic way. REFERENCES. Steel, A. J., Bruzzi, S. and Clarke, M. E., A 300 W 1.8 K refrigerator and distribution system for the CERN superconducting RF particule separator. International Cryogenic Engineering Conference 6 (1976) editeur, pages 2 Benda, V. and al, Cryogenic infrastructure for superfluid helium testing of LHC prototype superconducting magnets. Advances in Cryogenic Engineering, Vol 39A, 641-648 3 Quack, H. International Cryogenic Engineering Conference (1978) London 4 Spath, F. and al., A 2 kW He refrigerator for SC magnets tests down to 3.3 K, International Cryogenic Engineering Conference 14 (1992) 56-59 5 Gistau, G. M. The Tore Supra 300 W - 1.75 K refrigerator. Advances in Cryogenic Engineering, Vol 31 607-615 6 Fuerst, J. D., Design, construction and operation of a two cylinder reciprocating cold compressor. Advances in Cryogenic Engineering, Vol 37B 795-800 7 Mulder, J., The use of an expansion ejector in a 5 W refrigerator at 1.8 K. International Cryogenic Engineering Conference 4 (1972) 111-114 s Lebrun, Ph., Claudet, G. and Tavian, L., Development of large capacity refrigeration at 1.8 K for the Large Hadron Collider. Kryogenika, Praha April1996, 54 -59. 9 Dixon, S. L., Fluid mechanics. Pergamon Press. ~0 Gistau, G. M., Villard, J. C. and Turcat, F., Application range of Cryogenic Centrifugal Compressors. Advances in Cryogenic Engineering Vol 35B 1031-1037 ~ Schustr, P., Vins, M., Brunowsky, I. and Tucek, L. Helium low temperature compressor. Proceedings of Kryogenika April 96, 78-82 ~zProduct catalog of Barber Nichols ~3Fuerst, J. Selection of cold compressors for the Fermilab Tevatron. Advances in Cryogenic Engineering, Vol 39A 863-869 14 Bevins, B. S. and al, Automatic pumpdown of the 2K cold copmpressors for the CEBAF central helium liquefier. Advances in Cryogenic Engineering Vol 40 (to be published) ~5Kauschke, M., Haberstroh, C. and Quack, H. Safe and efficient operation of multistage compressors systems. Advances in Cryogenic Engineering Vol 40 (to be published) ~6 Guignard, J. Ph. Contribution h l'rtude de la stabilit6 de fonctionnement et de l'adaptation de charge des compresseurs centrifuges multi-&ages. Internal note CERN 1993. ~7 Brown, D. P., Schlafke, A. and Wu, K. C. Cycle design for the Isabelle helium refrigerator. Advances in Cryogenic Engineering Vol 27 501-508 ~s Gravil, B. 10 years of operation of the Tore Supra Cryogenic system. NIFS Symposium on Cryogenic systems, Mai 1996 Toki, Japan ~9Chronis, W. C. and al, Commissioning of the CHL refrigerator at CEBAF. Advances in Cryogenic Engineering Vol 40 (to be published)
A Cryogenic Axial-Centrifugal Compressor For Superfluid Helium Refrigeration
L. Decker, K. L6hlein, P. Schustr* M. Vine*, I. Brunovsk3~ § L. Tucek , Ph. Lebrun + and L. Tavian + Linde Kryotechnik, CH-8422 Pfungen, Switzerland * ATEKO,Resslova 13, CS-50010 Hradec Kr~ilov6, Czech Republic § PBS, CS-59512 Velk~i Bite~, Czech Republic + LHC Division, CERN, CH-1211 Geneva 23, Switzerland
CERN's new project, the Large Hadron Collider (LHC), will use superfluid helium as coolant for its high-field superconducting magnets and therefore require large capacity refrigeration at 1.8 K. This may only be achieved by subatmospheric compression of gaseous helium at cryogenic temperature. To stimulate development of this technology, CERN has procured from industry prototype Cold Compressor Units (CCU). This unit is based on a cryogenic axial-centrifugal compressor, running on ceramic ball bearings and driven by a variable-frequency electrical motor operating under low-pressure helium at ambient temperature. The machine has been commissioned and is now in operation. After describing basic constructional features of the compressor, we report on measured performance.
INTRODUCTION CERN's new project, the Large Hadron Collider (LHC) [1] now under construction, will use superfluid helium as coolant for its high-field superconducting magnets and therefore require large capacity refrigeration at 1.8 K. This may only be achieved by subatmospheric compression of gaseous helium at cryogenic temperature. To stimulate development of this technology, CERN has procured from Linde, with ATEKO and PBS as subcontractors, a prototype Cold Compressor Unit (CCU) with a nominal flowrate of 18 g/s @ 1 kPa inlet pressure, with a pressure ratio of 3 and with an isentropic efficiency better than 60 %. These specifications also meet the need for a booster stage enabling to triple the capacity of an existing warm pumping unit (WPU) which provides 1.8 K refrigeration to the CERN cryogenic test station, an upgrade described in a companion paper [2]. The compliance of the system allows to handle low flow continuously varying over the range from 6 to 18 g/s, with a pressure ratio of 1 to 3 and with an isentropic efficiency better than 0.5. For a given mass-flow, the gas inlet temperature may vary depending upon the number of test stations in operation. Table 1 summarizes the main specifications of the CCU. Table 1 Main specifications of CCU Capacity Helium flow-rate [g/s] Suction Pressure [kPa] Discharge pressure [kPa] Pressure ratio Helium suction temperature [K] Isentropic efficiency 195
Nominal
Low-flow
18
6 to 18
1
0.6
lto3 5.3 to 3.5 >0.5
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BASIC CONSTRUCTIONAL FEATURES General Layout The main constituent of the CCU is a cold box which houses all components at low temperature. On the top flange, the cold compressor is mounted with vertical axis. A cartridge design allows to remove the rotating part of the compressor without breaking the vacuum of the CCU enclosure. Two bayonet connections to the compressor inlet and outlet constitute the interface with a separate valve box which contains two insulation valves and a bypass valve. Figure 1 shows the general layout of the CCU.
Vacuum guard
Connection bayonet . ~ o l d compressor cartridge
i
~
Frequency converter
Cold box abinet
Figure 1 General layout of the CCU Cold compressor description [3] The axial-centrifugal unshrouded wheel with a diameter of 118 mm is fixed to the bearing shaft by a titanium tube giving reduced heat inleaks. Ceramic ball bearings and a variable-frequency electrical motor drive operate under low-pressure helium at ambient temperature. Motor drive, bearings and wheel compose the cartridge. The cold part constituted by the diffuser and volute remains inside the CCU cold box and is supported from the top flange by a thin-walled stainless steel tube. Instrumentation Efficiency assessment of the cold compressor requires accurate temperature and pressure measurements at inlet and outlet. Two calibrated germanium temperature sensors and two absolute pressure gauges have been chosen to fulfill this requirement. Other standard pressure and temperature measurements allow to monitor purge, cooldown, pumpdown and warmup. To avoid risk of air inleaks, all warm components which are working below atmospheric pressure are placed in a vacuum guard.
COMMISSIONING PROGRESS Commissioning was performed in two steps. During August 1995, pressure and leak tightness tests as well as piping integrity were checked. Cooldown and warmup were also performed. Initial problems of electrical breakdown in very-low pressure helium in the motor housing precluded performance tests. After improving electrical insulation, a second test campaign took place in November 1995 to complete commissioning.
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RESULTS Cooldown and warmup These two operations take a few hours each. Cooldown is produced by the helium flow pumped by the WPU in series with the CCU. During this phase, the two insulation valves are open, the drive speed is fixed to 100 Hz and the bypass valve controls the pressure difference between inlet and outlet of the compressor. For warmup, the CCU is completely stopped. A small flow of warm helium gas is injected at the CCU inlet and discharged at the outlet in a recovery line. Static heat inleaks Heat inleaks were assessed by inlet-outlet temperature difference measurement with the compressor stopped and for different flow rates, yielding a value of 50 W. Due to an increased thickness of the volute support tube during manufacturing (to avoid uncertainties from assembly and unexpected problems during tests), the heat load is 16 W higher and thus affects the global isentropic efficiency. Nominal and low-flow operation In steady-state operation, the compressor controls its suction pressure Pin by adjusting the motor drive speed N. The inlet pressure then stays stable within 10 Pa, better than the required value of 50 Pa. For a given mass-flow rh, the volumetric characteristic of the WPU fixes the CCU outlet pressure and consequently the pressure ratio of the cold compressor. Several measurements were done varying mass flow from 6 to 18 g/s and inlet temperature Tin from 3.5 K to 5.3 K. Table 2 summarized the main results and figure 2 shows the measured points on the calculated compression field. The operating conditions of turbomachines are expressed in terms of reduced mass-flow rhr and speed Nr, defined as follows:
m/ in in0
r / in0 VTin
Tin 0 Pin
with subscript 0 for design conditions
Table 2 Nominal and low-flow measurements Point
rh [g/s]
Tin [K]
Pin [kPa]
Tout [K]
Pressure ratio
Isentropic efficiency
N [Hz]
Nr [Hz]
fiar [g/s]
1 2 3 4 5 6 7 8 9
8.1 10.2 12.2 14.2 16 18 18 18.2 18
4.42 3.85 4.68 4.34 3.69 3.46 4.17 3.31 2.87
1 1 1 1 1 1 1 1 0.95
6.55 6.09 7.74 7.60 6.80 6.75 8.16 6.52 6.03
1.25 1.57 1.92 2.27 2.57 2.95 2.95 2.98 3.40
0.19 0.34 0.46 0.52 0.54 0.57 0.56 0.56 0.57
228 265 353 376 367 381 430 375 366
203 252 306 338 358 384 394 385 404
9.10 10.70 14.11 15.81 16.43 17.90 19.65 17.70 17.16
Isentropic efficiency The measured isentropic efficiency is slightly below specification. The measurement error, which depends on those on temperature and pressure, is estimated to +_ 0.02. The higher static heat inleaks to the cold compressor, which account for a loss of efficiency between 0.03 and 0.06, depending upon their exact location, may thus explain the difference to the specified efficiency. Possible corrective actions are being investigated by the manufacturer. At low flow, the relative proportion of the static heat loads in the loss of efficiency increases and below 14 g/s, it is no longer possible to obtain the specified value of 0.5. Stall line measurement By adjusting a bypass valve on the WPU, it was possible to run the cold compressor at constant pressure ratio with smaller mass flow. In this way, one can reach and explore the stall limit of the wheel. Figure 2 shows the good agreement between the calculated and measured stall limits.
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10
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Reduced mass-flow [g/s]
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Figure 2 CCU reduced compression field (parameter is reduced speed Nr). Endurance test In the course of the second test campaign, the CCU was operated in fully automatic mode, for a total of 50 hours without interruption. This test gives confidence in the future use of the CCU but more operating experience is required in order to qualify the long-term reliability of the bearings and motor drive. Identical motor and bearings installed in a test rig simulating operating conditions at Linde's have so far gone through more than 6000 hours of continuous operation without any problem.
CONCLUSION With the exception of a slightly lower than specified isentropic efficiency, the CCU supplied by Linde, ATEKO and PBS fulfills the technical requirements. Studies to improve its performance are under way. Moreover, additional tests are in progress to qualify long-term performance and reliability of the drive and bearing system.
ACKNOWLEDGMENTS The authors wish to acknowledge contributions of A. B6zaguet, S Claudet, D. Lavielle, and B. Vullierme to the preparation and execution of the cold compressor tests.
REFERENCES The LHC Study Group, The Large Hadron Collider, Conceptual Design, CERN Report AC/95-05(LHC) (1995). Benda, V., Dauvergne, J.P., Haug, F., Knoops, S., Lebrun, Ph., Sergo, V., Tavian, L., and Vullierme, B., Upgrade Of The CERN Cryogenic Station for Superfluid Helium Testing of Prototype LHC Superconducting Magnets, paper presented at this conference. Schustr, P., Vine, M., Brunovsk3~, I. and Tu'~ek, L., Helium Low Temperature compressor, Proc. Kryogenika 96, Prague (1996)
Upgrade of the CERN Cryogenic Station for Superfluid Helium Testing of Prototype LHC Superconducting Magnets
V. Benda, J.P. Dauvergne, F. Haug, S. Knoops, Ph. Lebrun, F. Momal, V. Sergo, L. Tavian and B. Vullierme LHC Division, CERN, CH-1211 Geneva 23, Switzerland
The cryogenic infrastructure of the station for testing LHC prototype superconducting magnets in superfluid helium below 2 K has been upgraded. Liquid nitrogen precooling has permitted to increase the liquefaction capacity of the refrigerator. The addition of cold centrifugal compressors with a pressure ratio of 3:1 has boosted the capacity of the warm pumping unit. To ensure adaptation of the pumping capacity, a heater-and-valve box allows to bypass the cold compressors. This box also comprises a 32 kW electrical heater for warming up the low-pressure gaseous helium before it enters the volumetric warm pumping unit. Possible impurities in the helium returning from the subatmospheric circuits are trapped in a freeze-out helium cleaner. Automatic process control and supervision permit unattended operation and optimal management of the helium inventory.
INTRODUCTION In preparation for CERN's new project, the Large Hadron Collider (LHC) [ 1], we have been operating since 1993 a cryogenic station [2] for testing prototype superconducting magnets in superfluid helium below 2 K. In view of the development of full-scale magnet tests, as well as the operation of a 50-m long prototype magnet string, we have now upgraded the cryogenic infrastructure of the test station.
LIQUID NITROGEN PRECOOLING A liquid nitrogen precooler was ordered to Air Liquide in 1995 and commissioned at CERN at the beginning of 1996. This device, connected to the 6 kW @ 4.5 K Air Liquide refrigerator, permits to increase the liquefaction capacity from 18 to 37 g/s. To reach this liquefaction capacity, a 1.85 MPa helium flow of 40 g/s coming from the high-pressure side of the refrigerator cycle, is precooled down to 81 K by means of a liquid nitrogen economizer, and reinjected at the corresponding temperature level in the refrigerator. As it is preecooled, this flow is purified in a 260 1 "Silicagel" adsorber. Figure 1 shows the flow-scheme of the precooling system, which consumes 700 1/h of liquid nitrogen. The adsorber is designed for 250-hour autonomy but the overall autonomy is still limited bythe main refrigerator adsorber. The impurity level is continuously analyzed at the bottom of the adsorber bed and a 3-hour regeneration cycle is initiated after impurity detection. The "Silicagel" bed is warmed up by gaseous nitrogen flow and cleaned by pumpdown and flushing with pure helium. The precooler is designed for a maximum mass flow of 80 g/s, which would yield a liquefaction capacity of 75 g/s. To reach this, a supplementary upgrade of the refrigerator is required. In particular, the expansion turbines have to be modified for higher flow and cycle compressors have to be added. Rapid cooldown and warmup of the superconducting magnet under test are performed by forced flow of gaseous helium under pressure, presently tapped from the cycle compressors of the refrigerator. 199
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EVACUATION GHe RECOVERY "SILICAGEL" A GN2 SUPPLY ANALYSER
LN2 SUPPLY PRECOOLED He, 81 K, 1.85 MPa (TO REFRIGERATOR) PURE HP He SUPPLY
Figure 1 Flow-scheme of the liquid nitrogen precooling unit
PUMPING CAPACITY UPGRADE Figure 2 shows the upgraded flow scheme of the pumping system which integrates two alternately used cold compressor units (CCU1 and 2), a heater-and-valve box (HVB), the existing warm pumping unit (WPU) and a freeze-out cleaner. HVB He REFRIGERATOR~ LP LINE
- I
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Figure 2 Flow-scheme of the pumping system Heater-and-valve box To ensure continuous adaptation of the pumping capacity over the range of flow-rate produced by the user devices, the HVB allows to bypass the low-pressure gaseous helium from the cold compressors. Below 6 g/s the WPU can handle the pumping flow alone; for higher flow, the coupling of one CCU is required. Two DN125 mm isolation valves, connection bayonets and U-tube transfer lines permit to connect alternately either CCU. For CCU commissioning, a dedicated helium test cryostat (HTC) is connected to the HVB to generate adjustable helium flow. This cryostat is also used for testing special 1.8 K components. The HVB also comprises a low-pressure drop, 32 kW electrical heater [3], for warming up the gaseous helium at 1 to 3 kPa before it enters the WPU. This heater is constituted of two 16 kW cartridges in series. Longitudinal copper plates, with a total exchange surface of 10 m 2, are heated by coaxial heating elements, brazed under vacuum to the plates for good thermal contact. To avoid electrical breakdown in very-low pressure gaseous helium, the coaxial heating elements are helium-tight and all electrical connections are outside the low-pressure vessel. Figure 3 shows the heater cross-section and Figure 4 the HVB. This heater has to handle large flow variations and due to its big thermal inertia, is controlled by the combination of a PID algorithm and an open loop calculating heating demand from measured flow. The heater is powered by a pulse-width modulation, 3-phase 400 V power converter.
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Cold compressor units The addition of cold centrifugal compressors with a pressure ratio of 3:1 has boosted the flow capacity of the warm volumetric pumping unit (WPU) from 6 to 18 g/s at 1 kPa suction, thus providing a useful refrigeration capacity of 360 W @ 1.8 K. The CCU supplied by Linde [4] is now commissioned, the other CCU supplied by Air Liquide has run and is being improved. Freeze-out helium cleaner Possible impurities in the 18 g/s helium flow returning from the subatmospheric circuits are trapped on-line in a freeze-out cleaner operating at atmospheric pressure, the design of which has been described in reference [5]. Figure 5 shows the flow-scheme of this purifier. The contamination is removed by cryotrapping sequentially in a freeze-out element and in a filter operating at a temperature of 30 K. Cooling of the gas is obtained by injection of liquid helium at a rate of 10 % of the main stream. GN2 SUPPLY GN2 VENT LINE GHe RECOVERY
/
/
REFRIGERATOR GHe HP LINE
r~ z
@
'L r
REFRIGERATOR GHe LP LINE
Figure 5 Flow-scheme of the freeze-out cleaner
LHe SUPPLY
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ICEC16/ICMC Proceedings
The system is designed to trap 1 kg of air before saturation, resulting to a minimum autonomy of about 300 hours. Regeneration is initiated when the purifier pressure drop increases up to 15 kPa. The system is warmed up to 90 K and purged with gaseous helium. Such a cleaning takes 30 minutes. High temperature regeneration, with nitrogen circulation, is not automatic and is only used in case of water contamination. During the regeneration, the pumping flow can be directly injected into the suction side of the cycle compressors if the impurity level is acceptable, or else is recovered in gas bags.
PROCESS CONTROL AND SUPERVISION All additions to the upgraded system are automatically controlled, thus allowing unattended operation. The liquid nitrogen precooler and the freeze-out cleaner are controlled by the existing ABB system of the refrigerator. The cold compressor units have their own Siemens programmable logic controller (PLC). A new "Utility" PLC controls the heater-and-valve box and the liquid helium distribution to the different test stations. An Ethernet| network interconnects the different PLCs, so that the "Utility" PLC is able to manage the helium inventory as well as the instantaneous consumption and availability of fluids and utilities for each test station. Operator interface is constituted by local operator panel and/or by workstations and X-terminals running a FactoryLink| supervision software. In total, 11 PLCs are simultaneously operating, while 3 workstations and 5 X-terminals manage the supervision of the test station area.
CONCLUSION With a total liquefaction capacity of 37 g/s and with a total pumping flow of 18 g/s at 1 kPa suction which provides a useful refrigeration capacity of 360 W @ 1.8 K and allows the parallel supply of 4 test stations, the cryogenic infrastructure of the test station has now reached the required capacity for testing full-scale magnets and the 50-m long prototype magnet string for the coming years. Comprehensive process control and supervision permits to operate the test stations with minimum staff. The next step will be to install a second 6 kW @ 4.5 K refrigerator available at CERN and to add dedicated circulation compressors for magnet cooldown and warmup.
ACKNOWLEDGMENTS We would like to acknowledge the contributions of our colleagues S. Claudet, P. Bernard, G. Bonfillou, G. Bochaton, A. Delattre, L. Herblin, J.P. Lamboy, D. Lavielle, H. Rieder, A. Tovar and A. Wiart of CERN, as well as M. Bonneton and D. Nuzzo of Air Liquide.
REFERENCES The LHC Study Group, The Large Hadron Collider, Conceptual Design, CERN Report AC/95-05(LHC) (1995). Benda, V., Duraffour, G., Guiard-Marigny, A., Lebrun, Ph., Momal, F., Saban, R., Sergo, V., Tavian, L. and Vullierme, B., Cryogenic Infrastructure for Superfluid Helium Testing of LHC Prototype Superconducting Magnets, paper presented at CEC, Albuquerque (1993). Benda, V., Sergo, V. and Vullierme, B., Electrical Heater for very-low pressure helium gas, Proc. Kryogenika 96, Prague (1996) Decker, L., L6hlein, K., Schustr, P., Vine, M., Brunovsk3~, I., Tu~ek, L., Lebrun, Ph. and Tavian, L., A Cryogenic Axial-Centrifugal Compressor For Superfluid Helium Refrigeration, paper presented at this conference. Dauvergne, J.P., Delikaris, D., Haug, F. and Knoops, S., A Helium Freeze-out Cleaner Operating at Atmospheric Pressure, paper presented at CEC, Colombus (1995).
Performance Analysis of Multistage 80K Centrifugal Compressors for Helium Refrigerator
Hiroshi Asakura,* Nobuyoshi Saji,* Yukio Kaneko,* Shoichiro Yoshinaga,* Mikio Mori,* Junichi Sato,* Akihiro Miyake,* Tsutomu Iwasaki,* Izumi Nishimura,* Takashi Hosoya,* and Tomohiro Umeda** *General Machinery Division, Ishikawajima-Harima Heavy Industries Co.,Ltd., 3-2-16,Toyosu, Kotoku, Tokyo, 135 Japan **Super-GM, 5-14-10, Nishitenma, Kita-ku, Osaka, 530 Japan
A completely oil-free centrifugal compressor system which has 4 stages for advanced helium refrigerator has been developed under the Japanese national project 'SuperGM.' A high performance of each stage compressor was confirmed by the aerodynamic performance test at 80K. The multistage performance test has been carried out for the purpose of studying the flow adjustment method to control the liquefaction rate. With the multistage performance data, the simulation analysis of refrigeration system has been done, and appropriate controllability of liquefaction rate by varying the 1st stage compressor speed has been confirmed.
INTRODUCTION An advanced type helium refrigeration system for a superconducting generator is now being developed in Japanese national project 'Super-GM.' The features of this refrigeration system are high reliability which is attained by adopting complete oil free cold compressors, and simple controllability of liquefaction rate by varying compressor speed. We have already developed single cold compressor unit which is the key component of this system, and confirmed that it can meet with each specification sufficiently both in mechanical and in aerodynamic performance. [ 1,2] In this paper, we will report the multistage characteristic attained by combined compressor performance test and simulation analysis of the refrigeration system when the compressor speed is varied.
ADVANCED TURBO TYPE REFRIGERATION SYSTEM General Features The flow sheet of the refrigeration system is shown in Figure 1. This system is so designed as to provide a liquid helium of 120 L/h with two expansion turbines and a J-T valve. Helium gas pressure is raised by 4 compressors from 0.1 MPa to 0.8 MPa. The gas from the high temperature expansion turbine returns to the inlet of the 2nd stage compressor. The purpose is to prevent efficiency drop caused by too small volumetric flow in the latter stage compressors. Since each compressor is independent, and its speed can be varied with the inverter, the liquefaction rate of the refrigeration system can be varied by changing compressor speed with constant J-T valve opening. Oil Free Centrifugal Cold Compressor Each stage compressor has a common structure that a vertical rotor with a impeller at bottom end is rotated 203
204
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by an electric motor located in the center and all of them are hermetically sealed in one casing. Magnetic bearing which is employed as an oil free bearing contributes to reduce the vibration which may be caused by a distorted shaft at low temperature, thanks to its automatic balancing system. The electric motor can get a high power density of 25 kW at 100,000 rpm by adopting a claw-pole type synchronous motor excited externally.[ 1] The impeller is of a shroudless and high efficiency three dimensional blade, and is made from high strength aluminum alloy. The maximum peripheral speed reaches 560 m/s that contributes to increase pressure ratio. The aerodynamic performance of each stage compressor was tested in a closed loop of low temperature helium gas and was confirmed to exceed the specific point sufficiently.J2]
LIQUEFACTION RATE CONTROL BY VARYING COMPRESSOR SPEED Multistage Performance of The Compressors We carried out tests by combining each two of the four compressors to catch characteristics extending the whole flow range of the compressor system with intermediate flow between the 1st and the 2nd stage. Since the surging that is unstable phenomenon at the low-flow rate range of the turbocompressor depends on the characteristics of the circulating piping system, the consideration was paid so that the combined compressor test apparatus gave equivalent circulating volume to the actual refrigeration system. Though the speed of each stage can be changed independently, it is efficient to keep the pressure
ICEC16/ICMC Proceedings
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ratio for the high temperature expansion turbine. Therefore, characteristics of the compressors were tested by varying the 1st stage compressor speed with the 2nd through the 4th stage compressor speed kept at each rated speed. Figure 2 shows the test results in the form of multistage characteristics of 4 stages. Figure 3 shows isothermal efficiency in the form of 4 stage compressor system.
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Refrigeration System Simulation We constructed a simulation program of the refrigeration system including the multistage characteristics of the compressor system to examine how liquefaction rate and compressor power consumption vary when the compressor speed varies. The characteristics of each stage compressor was described in non-dimensional function based on the test results. Opening of the J-T valve was kept constant. Figure 4 shows analyzing result with simulation, that is change of liquefaction rate and compressor power consumption per unit liquefaction rate which stands for the figure of merit in this system when the 1st stage compressor speed was changed. As shown in the figure, the liquefaction rate can be controlled over the range of 50 - 160 L/h only by changing the 1st stage compressor speed in the range of 40,000 - 100,000 rpm. And it is found that liquefaction rate can be varied in the range of 100 - 160 L/h with almost constant power per liquefaction rate. 0.70
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ICEC16/ICMC Proceedings
When the system balances after compressor speed was changed, the operating point of each stage compressor is automatically fixed. The black points in Figure 2 and 3 show the operating points at each speed. As shown in Figures, it is confirmed that each operating point is in the stable range sufficiently apart from the surge line and that isothermal efficiency is almost at the peak point.
CONCLUSION We carried out performance tests by combining oil free cold compressors and clarified multistage compressor characteristics when the 1st stage compressor speed was varied. We made a simulation program of the refrigeration system including the compressor characteristics and analyzed variation of liquefaction rate and compressor power consumption when the speed was varied. As a result, it was found that liquefaction rate can be controlled over the range of 40 - 130 % of the rated liquefaction rate by varying the 1st stage compressor speed and that compressor power per liquefaction rate did not vary so much at that time. It was also confirmed that the operating point of each stage compressor is in the stable range sufficiently apart from the surge line.
ACKNOWLEDGMENT This research has been carried out as a part of R&D on superconducting technology for electric power apparatus under the New Sunshine Project of AIST, MITI, being consigned by NEDO.
REFERENCES 1 2
Asakura,H. et al. 80K centrifugal compressor for helium refrigeration system, In: Advances in Cryogenic Engineering 37B, Plenum Press, New York (1992) 787-794 Asakura,H. et al. Performance test results of 80K centrifugal compressor for helium refrigerator, In : Advances in Cryogenic Engineering 39A, Plenum Press, New York (1994) 893-900
Design, Manufacture and Consideration for Test Result of Centrifugal Cold Compressor for TEVATRON Lower Temperature Upgrade
Nobuyoshi Saji, Yukio Kaneko, Hiroshi Asakura Development Department,General Machinery Division,Ishikawajima-Harima Industries Co.Ltd.,2-16 Toyosu 3-chome, Koto-ku, Tokyo 135 Japan
Ten years have passed since Tevatron first operation. In 1994, Fermilab completed to grade up the refrigeration system including capable of lowering two-phase temperature down to 3.5K with the satellite refrigerators equipped with cold compressors. Since we supplied cold helium compressors, we will report the process of upgrade, structure of the cold compressor and operating situation up to date.
INTRODUCTION In April 1994, Fermilab announced that it discovered top Quark. Also it was announced in February 1996 that the Japan, U.S. and Europe research team acquired the data that smaller particles constituting Quark possibly exist. In Tevatron, about 800 superconducting magnets were arranged and forcedly cooled by supercritical helium (SHE) in the underground tunnel of 6km circumference. The refrigeration system for Tevatron used since 1984 is composed of 4000 L/h central helium liquifire and 24 satellites refrigerator. A satellite refrigerator cools each left and right 125m strings of superconducting magnets. JT valve located at the remotest place from the cold compressor as shown in Fig. 1. If the cold compressors should not be operated, the temperature of helium would be 4.54K at the two-phase dewar and approx. 4.9K at 125m upstream point of the JT valve outlet, showing approx. 400 mK higher temperature because of pressure loss of two-phase helium. Limited acceleration energy for the dipole magnet to quench at this temperature is 900
I SATELLITE REFRIGERATOR} COLD BOX
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125m
MAGNET STRING
Fig 1 Satellite Retrigerator with Cold Compressor
_1
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Fig 2 Appearance of Tevatron Cold Compressor 207
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Gev. The energy can be raised up to 1050 Gev when liquid helium is depressurized by the cold compressor and helium temperature is lowered down to 3.56K at the entrance of the compressor. [1 ] However, when cold compressors are installed, the driving power plus heat leak through the shaft and casing become thermal load to the refrigerator. Therefore, refrigeration capacity was increased to meet with the power of 24 compressors.
COLD COMPRESSOR The apparatus to depressurize liquid helium by directly inhaling evaporated gas and lower its temperature is the cold compressor. Its appearance is shown in Fig.2. The specification of the cold helium compressors installed in Fermilab is shown in Table 1. The crossection of the machine is shown in Fig. 3. The shaft is supported by the dynamic gas bearings in the complete oilfree structure. The motor is air-cooled by a fan to eliminate possible trouble in the cooling water system and to get better maintenance. The cold section is in the vacuum chamber and both the inlet and outlet pipes are so made to be vacuum multi-layer insulation. The very small impeller made with aluminum alloy precision casting was proved, by Fermirab who operated a prototype delivered in 1990, to attain a high performance in the wide flow range. [2] The impeller inhaling cold helium and the motor driven at room temperature are placed in the distance and the shaft was of a hollowed shape for Insulation. The periphery of the aluminum die casting endrings placed at the both ends of the motor rotor was enforced by titanic alloy to produce a high speed motor with the maximum rotating speed of 95,000 rpm. A thin cylinder was used to connect between the compressor casing and drive section casing. Heat capacity coming into the compressing section from the driving unit throughthe shaft and casing is 30 W. However, its influence to efficiency remained approx. 3 %. Dynamic Gas Bearing The journal bearing is a foil type dynamic gas bearing of a simple structure made by triple-coiled 0.04 mm thick stainless sheets and the thrust bearing is a spiral groove type. The section which contacts the shaft at start and stop is coated with Teflon. Satellite refrigerator often inhales liquid helium along with evaporated gas into the cold compressor. Since this happening is difficult to prevent, Fermilab made the several times test to have liquid helium inhaled into the fast revolving impeller which used prototype cold compressor. The compressor stopped without noise and could be started again immediately after reset the invertor. Coo
Table 1 Specification of Tevatron cold compressor Inlet pressure Inlet temperature Outlet pressure Mass flow rate Rotating speed Operation range Motor power Impeller diameter Type of Bearing
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--
- .....
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0.5 atm (min. 0.4 atm) 3.56 K (min. 3.38 K) 1.4 atm 60 g/s (range 40 to 70 g/s) 80000 rpm (max. 95000 rpm) 20000 to 95000 rpm 1.4 kW (maximum) 33 mm Dynamic gas bearing
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Cryostat The cryostat is designed with the utmost attention to compactness. The inlet and outlet pipes are double pipes each with an external diameter of 35mm. The compressor is connected with bayonet joints to the valve box rising, on a 220 mm-diameter pipe, from the strings of superconducting magnets located in the 6.5 m underground. The valve box is of a stainless make with a diameter of 800 mm and a height of 1,500 mm equipped with a heat exchanger with a capacity of 130 litters liquid helium and ten or more valves or ports. Since these equipment is arranged on the 6 m pipe, the structure is unstable. Influence of vibration of the valve box to the cold compressor was anticipated from the planning. The root of the valve box is secured with three turnbuckles, but vibration far larger than the permissible value of l~tm prescribed by us was felt by hand. OPERATION The 27 cold compressors delivered in autumn 1992 were operated in December 1993, built in the whole 24 satellite refrigerators of Tevatron. Several problems occurred with operations made before and after the incorporation. At the first operation the rotor was forced to stop one day after the start, caused by impure gas from the motor. [4] Fermilab baked the motor section at 330 K for 24 hours before operation. However, the problem should have been solved by baking at 100 ~ for approx. 100 hours. The electric insulation of the motor can resist up to 125 ~ The rotating speed of the cold compressor is controlled by the computer which keeps pressure at the compressor inlet or two-phase helium dewar at constant. When the set of the temperature of two-phase helium in the magnets is changed from low to high temperature, helium in the magnets temporarily subcooling and has come not to evaporate. The level of liquid helium in the dewar rises, liquid helium overflows and the cold compressor trips by inhaling its. As countermeasures, enlarging the capacity of the two-phase dewar or changing temperature slowly are considered. However, Fermilab took the method to install a heater in the two-phase dewar. Trips of the Cold Compressor at the Times of Quench of the Magnets As an incident of magnet quench, the outlets of the cold compressors are connected by the common pipe and the motor tripped at the opposite place of Tevatron ring approx. 3km apart from the magnet which caused quench. In the compressor operated at 42,000 rpm, outlet pressure first increased and inlet pressure also increased 4 seconds later. Further 5 seconds later compressor rotating speed began to rise, reaching 80,000 rpm in 15 seconds. Further 15 seconds later, the motor stalled and tripped. In this instance, the flow rate of evaporated gas from the two-phase dewar decreased when inlet pressure began to increase and the compressor entered low-flow rate operation. Since the compressor was in slow speed operation, apparent surge phenomenon did not occur. However, since rotating speed was sharply increased, the compressor entered a strong surge, inhaled liquid helium and tripped. After the trip, rpm did not change because frequency generated by the invertor was plotted. The rotor remained at stop. began to lower, the computer gave a direction to lower speed to 80,000 rpm. Since rotating speed was reduced in haste, inlet pressure (dewar pressure) increased suddenly, evaporation volume decreased and the compressor entered surge. Each of the above cases was solved by making the control inch by inch and slowly and effect the heater in the two-phase dewar. However, it is difficult to explain why the motor tripped frequentry at the opposite point of the ring 3 km apart from the quenched magnet. Lately Operation After the cold compressor was operated for 28,626 hours by August 1994, the journal bearings were changed to bearings possible for low-speed operation at standby time to reduce refrigerator load and the minimum speed was changed from 40,000 rpm to 20,000 rpm. The used bearings were in good conditions. After that, the compressor was operated for 93,000 hours by April 1995. [5] During the time after the minimum speed was lowered to 20,000 rpm, bearing failure occurred three times. The possible cause is that by reducing rotating speed, load capacity of the dynamic pressure gas bearings decreased and the shaft and bearings slightly contacted when the valve box vibrated. Some countermeasure to stop vibration is required. How-
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ever, after the minimum speed was raised to 30,000 rpm temporarily, no failure has been reported yet so far. According to Felmilab's report in April 1996, total operating time of 24 satellites was 130,249 hours and average operating time of one satellite was 5,427 hours. Efficiency of the Cold Compressor The adiabatic efficiency measured during the operation in autumn 1993 was a very low value of 40- 50 %. The designed efficiency of 70 % at the design point and 60 % or more at the working area had been cleared with allowance even with the prototype operation. [2] The difference between the prototype and 27 compressors lies in adoption of aircooling for the latter motors and inlet and outlet pipes for the cold compressors. One of the causes is that Fermilab inserted a suction filter in the inlet pipe to prevent recurrence of the accident that cuttings were inhaled during prototype operation. The adiabatic efficiency increased by 10 15 % by removing the filter. [3] Table 2 shows the difference of pipe bore between the prototype and 27 compressors concerning the inlet and outlet pipes. The reason of the change was based on Fermilab's request for downsizing with light weight of the cryostat including bending sections. According to calculation, pressure loss at the pipe sections reaches 13 % of the compressor head at the maximum flow rate. If the pipe size is made similar to that of the prototype, efficiency will be increased by 10 %. Table 2 Comparision of inlet and outlet pipe bore between prototype and 27 machines Inlet pipe
Outlet pipe
Prototype
mm
30.7
23.9
27 machines
mm
20.45
17.5
2.25
1.87
Ratio of velocity (1.0 for prototype)
CONCLUSION The cold compressors for Tevatron have been confirmed their reliability for a long operation and become practical machines that can be operated in great numbers at the same time. However, the cryostat vibration must be solved.
ACKNOWLEDGMENT We express our thankfulness to the fact that Fermilab decided to choose cold compressors for Tevatron satellite refrigerators and persistently continued operations despite several troubles. Also we appreciate their kindness to provide some data for this paper.
REFERENCES 1 2 3 4 5
Fuerst, J. In:Advances in Cryogenic Engineering Vol.35 Plenum Press, New York (1990)1023-1030 Fuerst, J. In:Advances in Cryogenic Engineering Vol.39 Plenum Press, New York (1994)863-870 Theilacker, J., Tevatron Cold Compressor Operating Experience Criogenics (1994) 34 ICEC supplement 107-110 Norris, B. L.,Status Report on the Tevatron Lower Temperature Upgrade Criogenics (1994) 34 ICEC supplement 73-76 Norris, B. L., Initial Performance of Upgraded Tevatron Cryogenic Systems (WE-BE-6)
Development of Helium Oil Free Screw Compressor
K. Kitagawa, Y. Hirao, Y. Yanagi and Y. Ikeda* Mayekawa Mfg. Co. Ltd., Aza-Okubo, Moriya-Machi, Kita-Sooma-gun, Ibaraki 302-01, Japan *Super-GM,5-14-10, Nishitenma, Kita-ku, Osaka 530, Japan
It is well known that impurities originated in oil lubricants of helium compressors markedly affect on the reliability of refrigeration system. Hence Mycom has been concentrated on development of an oil free screw compressor since FY 1987. This paper will represent performance data of the oil free screw compressor which was comprised of non contact tapered rotors, taking into thermal deformation applying for high rotational beatings and seal systems based on R&D of an elementaw technology. In addition we demonstrate that it can be attained an isothermal efficiency of 50 % under a throughput of 125 g / s , an outlet pressure of 0.8 MPa and an inlet pressure of 0.1 MPa and a suction temperature of 250 K.
INTRODUCTION Super-GM ( Engineering Association for superconducting Generation Equipment and Materials ) was established in September, 1987 to develop technology for superconducting generation equipment and materials as a national project. Since then we have been developing two types of high reliable refrigeration system lbr SCG ; one is a conventional refrigeration system for a 70 MW class SCG using oil injected screw compressors, and the other is an advanced refrigeration system for a 200 MW pilot machine using oil free compressors which perfectly prevents from the troubles due to impurities. MYCOM has much concentrated to develop an oil free screw compressor which compresses helium gas over retaining a clearance inbetween screw rotors under a high rotational speed. So that the main subjects are to develop technology, for non contact tapered rotors taking into a thermal deformation, beating and seal systems for a high speed rotation as 25,000 rpm.. This paper represents a planning of test machine using air, pertbnnance testing and operational control. Furthermore, based on the results of test machine we discussed planning of a trial helium compressor. DEVELOPMENT OF ELEMETAL TECHNIQUE A test stand using air was planned to obtain the per~brmance of main components as beatings, seals and rotors under the high rotational speed. Figure 1 shows the cut out view of the test stand, where timing gears was equipped at the end of the shaft to maintain a non contact condition, with ball and roller bearings suitable for a high speed operation, and two staged seals composed of an oil and a gas seals were provided to prevent lubricant oil of the beating from penetrating into the rotors. In addition tapered rotors were applied to ease the thermal deformation due to a temperature difference between inlet and outlet temperatures. The test rotor diameter was 102 mm. .High Speed Bearings For the oil screw compressor ball and roller beatings with a very high accuracy of positioning were chosen in order to maintain an infinitesimal rotor-lobe clearance in operation. A cylindrical roller beatings were applied for radial beatings and 4 point contact ball beatings for thrust beating. Especially for the thrust beatings a wpe of low pre-load was employed to keep an adequate contact angle under a high speed operation. The roller bearing has a high accuracy of positioning while it has not enough damping capacity of 211
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vibration to sustain a rigidity, so that it must be operated in the region far from the shaft critical speed. Its performance testing resulted in less than 5 u m of shaft vibration up to a 25,000 rpm and a stable operation and the thrust beating for M rotor was very stable until a maximum speed, whereas that of F rotor could be operated stably by setting a rather high pre-load. Gas seal 0il seal / /~ Radial b e a r i n g T h r u s t bearingi-(pTe -- load) \, , , Speed up R e a r \ [-~ ~ / Drive -shiaf--L-t\, ~ - ~ ~--~~3
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Figure 1 Cut out view of the test oil free screw compressor Gas Seal For the gas seal we tested a labyrinth, a honeycomb and an aerodynamic pressure seals as a non contact seal. An aerodynamic pressure seal was chosen to prevent the gas leakage into the rotor casing from the intermediate chamber, since it gave us a lower temperature of the intermediate chamber than that in using a honeycomb seal. Figure 2 shows a photo of the used aerodynamic seal. Oil seal For the oil seal a magnetic fluid seal of contact type was applied in order to seal tight which is useful for vacuum and high speed. The temperature of magnetic fluid seal in the seal region was found to greatly affect the seal pressure toughness. Hence the cooling method was improved by modifying a fin on the cooling side. It was confirmed that its seal toughness was 0.05 MPa under the conditions of a speed of 25,000 rpm and a compression ratio of 3.0. Figure 3 shows a photo of the used magnetic fluid seal.
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Crossover occurs with a probability Pc (typically, Pc e [0.6,0.8]). The mutation operator swaps the value of a bit (its value is changed from 0 to 1, or vice versa) with a probability Pm (typically, Pm C[0.001,0.01]). Usually, mutation has a negative effect on an individual, but occasionally it can change a poorly performing individual into a fit individual. Once the new generation of individuals is complete, the fitness of each individual is evaluated and a new generation formed, again based on the fitness of the parent individuals of the current generation. Usually, the algorithms terminated after a predetermined number of generations. OPTIMIZING THE DESIGN VARIABLES BY GA In our discussion, the objective function is chosen to maximize the isentropic efficiency of turboexpander subject to a set of constraints. It can be expressed as max r/s (p, ~t ) s.t. a set of constraints
(1)
where the optimization variablesare reaction p and diameter ratio/z. In 900 inward flow radial turbine the following relationship exists between the velocity ratio h-1and reaction p u-~ = ~/(1- p)~o: cos 2 a ,
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absolute Mach number at nozzle outlet less than 1.2. difference between flow angle at rotor inlet fll and blade angle at rotor inlet fl~ less than 3 ~
9 9
relative Mach number at rotor inlet less than 0.5. relative velocity at rotor outlet WE greater than relative velocity at rotor inlet Wl.
ICEC 16/ICMC Proceedings
223
l
flow angle at rotor outlet ranging in 800 - 100 ~ The working procedure of GA to optimize the design variablesp and/~ is shown as figure 1. During the process, the isentropic efficiency r/s is calculated by one-dimension flow method[2]. In the cases that calculations from the chosen individual of /9 and/~ do not satisfy the above constraints, we simply let qs be a small value, such as r/s =0.1%, then the corresponding individual in GA will have a low probability to reproduce.
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Figure 6 shows the time-base records of the rotor vibrations for the bearing(C=20, n--22) at high rotational speed 152,48 lrpnl The rotor orbit is clear and small. There is no obvious subffequency whirl occurred and the rotation is stable. The waterfall diagram of the rotor vibration for the bearing(C=20, n=22) is shown in figure 7. The major resonant ~equency was synchronous at every rotor speed. It had a maximum amplitude near the natural speed of the rotor(150,000rpm). However, the sub synchronous resonance is not significant and the frequency is low (about 200Hz). It shows the foil bearing with optimum parameters has high whirl stability. Durability Test The durability test in which the stag-stop cycle (up to 120,000rpm) was repeated 60 times were carried out. From the test results shown in figure 8, it is seen that the rotation characteristics did not change at all and the rotor stably rotated during the tests.
236
ICEC16/ICMC Proceedings L/D=1.0
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Figure 7. Comparison of observed velocity oscillations with calculation.
SUMMARY The oscillating flow in the pulse tube has been visualized by using the smoke-wire technique and the change of the velocity profiles during one cycle has been investigated. Typical velocity profiles, which are observed in oscillating viscous flows, are observed in the beginning of the compression and expansion processes. The velocity and displacement in the orifice pulse robe are larger than that in the basic one. The velocity oscillation is related to the differential of pressure variation with respect to time for the basic pulse tube and to both the differential of pressure variation and pressure difference between the pulse tube and reservoir for the orifice pulse tube. REFERENCES 1 Lee, J. M., Kittel, P. Timmerhaus, K. D. and Radebough, R., Flow patterns intrinsic to the pulse tube refrigerator, Proceedings of the 7th International Cryocooler Conference, PL-CP-93-1001, Santa Fe(1993) 125-139.
Intrinsic Behaviour of a Four Valve Pulse Tube Refrigerator
M.Tht~rk, H.Brehm, J.Gerster, G.Kaiser, R.Wagner, and P.Seidel Institut ft~r Festk6rperphysik, Friedrich Schiller Universitfit Jena, Lessingstr.8, D-07743 Jena
The oscillating behaviour of working gas inside a single stage four valve pulse tube refrigerator (FVPTR) with co-axially arranged regenerator around the pulse tube has been investigated. Pressure and temperature oscillations have been measured at various points. A wall temperature distribution along the regenerator has also been obtained. The oscillation behaviour of the gas pressure is quite similar at all positions through the pulse tube. However, the one of the gas temperature changes with the location along the pulse tube dramaticly. The amplitude of the temperature oscillation near the center of the pulse tube is larger than that at the ends. The phase of temperature oscillations in comparison to the pressure oscillations shifts over the length of the pulse tube. The observed datas of the temperature and pressure behaviour were used to identify typical loss mechanism of the FVPTR.
INTRODUCTION The pulse tube refrigerator containing no moving parts in the cold section is more attractive for higher reliability, simpler construction and lower vibrations of mechanical and electromechanical origin than any other regenerator typ refrigerator. In the last years, the refrigeration power, and the lowest achievable temperature have been improved by the modification from the original basic pulse tube [1 ] over the orifice pulse tube [2] to double inlet pulse tube [3], multi-bypass pulse tube [4], and four valve pulse tube [5,6]. Recently, the lowest temperature that has ever been reached by a pulse tube was realized near the lambda point of helium [7]. Compared with the improvement of performance, a clear explanation of basic phenomenas inside of a pulse tube refrigerator has not been given even though several analytical models have been presented [8,9]. On the other hand, a few experimental studies [10] have been reported, in which the intrinsic behaviour of the orifice pulse tube refrigerators was investigated systematically. Neverthless, open questions about the special intrinsic effects inside the FVPTR remained. In order to understand the internal processes occurring in the FVPTR this paper will show some test results. And the main objective is revealing the effects of the oscillating flow phenomena of working gas inside the pulse tube on the thermal behaviour of a FVPTR in order to gain a better understanding of the real working conditions. The results shall also be used to obtain practical rules for system sizing and to provide datas for testing our theoretical model [ 11 ].
EXPERIMENTAL SET-UP The tested pulse tube and regenerator are in practical co-axial configuration. A sketch of the experimental apparatus is shown in fig. 1. The main datas of the construction of the FVPTR are the same as reported in reference [6]. Measurement methods have been used which allow instantaneous measurements of mass flow rate, pressure, and temperature in a FVPTR during the actual operation. Piezoelectric pressure sensors are used to monitor the pressure oscillations at the hot ends of the regenerator, pulse tube as well as in front of the rotary valve. As indicated in fig.l, measurements of the gas oscillations were made within the pulse tube at 13 points using resistance wire sensors (RWS). The RWS were constructed out of tungsten wire of 10mm lengths and with a diameter of 5gm and are suspended in the fluid flow by the pulse tube wall. The response time of the pressure transductor and the RWS is about 3gs (300 kHz) and - T } m Tdt where cp is the heat capacity at constant pressure, T is the temperature, m the mass flow, and I: the time it takes to complete one cycle. If the graph of the mass flow and the pressure oscillation in the FVPTR comes close to the rectangular shape as it is shown in figure 2, the enthalpy flow come easily down to (H) - Cp. m[Tout- Tin]. Increasing the refrigeration power is possible by a larger amount of exchanged working gas and/or a larger temperature difference. For practical reasons the input temperature T,, is fixed by the ambient temperature. Therefore only that amount of outflowing mass exchanged at the hot end which has a higher temperature than the ambient temperature increases (H). And now it is important to know how the temperature alters at each point within the pulse tube and how it depends on the different working conditions. A typical temperature behaviour of the working gas along the pulse tube length within the time intervall of compression and expansion, respectively is documented in figure 3. Figures 4 and 5 show the effects of some important working parameters on the temperature behaviour inside the pulse tube. There we are able to see that if the opening of the needle valve (OVN) is enlarged the mass flow increases as well, but the temperature in the center of the pulse tube increases during a short period of time up to 450 K. This results in a high conjugated heat transfer between pulse tube and regenerator and high axial heat loss, respectively. The high temperature rise at the cold end during the inflow periode indicates us a large blow down loss.
ICEC16/ICMC Proceedings
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262
ICEC16/ICMC Proceedings 400
300
200
100'
0 0
0.2
0.4
0.6
0.8
Time [ s I
Figure 5" Temperature and pressure oscillations at different locations in a pulse tube Typical Lissajous figures of the temperature vs pressure in the pulse tube were shown in fig 6. These figures show the temperature and pressure of the gas oscillating near the center of the tube, but at different positions along the tube. It is observed clearly that the shapes depend on the location in the pulse tube. The shape of these Lissajous figures are very sensitive to opening of the needle valve. HOT END
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Figure 6" Lissajous figures temperature versus pressure at hot end, center, and cold end CONCLUSIONS Measurements of instantaneous mass flow rates and the temperature have been made on a FVPTR during actual refrigerator operation. These measurements have allowed us to determine the gross refrigeration power experimantally as well as phase angle relationships between oscillating parameters in this system.The consequences of the working parameters on the temperature behaviour within the pulse tube is of great importance for the further improvement of the FVPTR. Further systematic investigations on this are currently under way. This work is supported by the German Ministry of Science and Technology under contract no. 13 N 6618 REFERENCES
7 8 9 10 11
Gifford,W.E. and Longsworth,R.C.,Trans.ASME J. Eng. Ind (1964) 63264. Mikulin,E.I., Tarasov,A.A. and Shkrebyonock,M.P., Adv. Cryo. Eng. (1984) 2__29629. Zhu Shaowei, Wu Peiyi, and Chen Zhongqi, Cryogenics (1990) 30514 Cai, J.H., Wang, J.J., Zhu, W.X., and Zhou,Y.,Cryogenics (1994)34713 Matsubara, Y.,Tanida, K.,Gao, J.L.,Hiresaki, Y.,and Kaneko, M., Proc.Fourth JSJS, Beijing, PR China (1993) 54 J.Blaurock, R. Hackenberger, P.Seidel and M. Thtirk, Cryocoolers 8, Ed.R.G.Ross, Jr., Plenum Press, New York(1995) 395 Thummes,G., Bender, S.,and Heiden,C., Cryogenics (1996) 3.__66,will be published Radebaugh, R.,Adv. Cryo. Eng. (1990) 3__551191. Tominaga,A.,.Proc.Fourth JSJS, Beijing, PR China (1993) 79 Rawlings,W., Radebaugh, R., and Bradley, P.E.,Adv. Cryo. Eng, (1994) 39B 1149. Kaiser,G., Brehm,H., Thtirk,M., and Seidel,P., Cryogenics (1996) 3__66,accepted
Anomaly of One-Stage Double-inlet Pulse Tube Refrigerator
Toyoichiro Shigi, Yoshiaki Fujii, Masahiro Yamamoto, Masaki Nakamura, Minoru Yamaguchi, Yoshiko Fujii, Tomio Nishitani*, Tetsuya Araki*, Etsuji Kawaguchi**, Masayoshi Yanai** Dept. of Applied Phys.,Okayama Univ. of Science, Ridai-cho 1-1, Okayama 700, Japan * Shiga Technology Center, Iwatani Intern. Corp., Katsube-cho 1095, Moriyama 524, Japan ** Iwatani Plantech Co., Ltd., Katsube-cho 1095, Moriyama 524, Japan
An one-stage double-inlet pulse tube refrigerator generated a low temperature 23.9 K.
Anomalous temperature variation occurred in the pulse tube against the bypass
valve opening, which is considered to be common in this type of the pulse tube refrigerator.
In order to make clear the reason for this anomaly, experiments were
performed and reasonable conclusion was obtained.
INTRODUCTION Among many modifications of the pulse tube refrigerator, the double-inlet type devised by Zhu et al. in 1990 [1] has great advantages in its improved performance and its simplicity.
So, we decided to
investigate the performance of an one-stage double-inlet pulse tube refrigerator in detail. CONSTITUTION OF THE REFRIGERATOR This refrigerator consists of a compressor ( suction volume capacity 4 m3/hr, max. pressure 2.1 MPa and power consumption 2.2 kW ), a rotary valve which generates nearly sinusoidal pressure variation ( max. pressure 2.1 MPa, min. pressure 1.2 MPa ), a regenerator containing about 2000 brass screens ( 200 mesh ) in a stainless steel tube ( ID 38.0 mm, length 21 cm ), a Cu-Ni pulse tube ( ID 17.4 mm, length 42--36 cm ), a gas reservoir of 1g, an adjustable orifice valveand an adjustable bypass valve. EXPERIMENTAL
[1] At first, the refrigerator was operated at the optimum condition with the pulse tube 40 cm long.
The
lowest temperature 23.9 K was obtained at the condition : frequency 1.72 Hz, orifice valve opening Cv 0.0137 and bypass valve opening Cv 0.0228. At the same condition, the cooling power of this refrigerator
was measured as shown in Figure 1. This operating condition corresponds to the point A in Figure 3. 263
Of course, the shape of the curve is
264
ICEC16/ICMC Proceedings T6
(w) Q
T5 2
-
Tll .
~,~.~
360ram %
30 T (K)
50mm
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,9 [
T9 + 50ram
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T3
74ram
0ram
T2 74ram
55mm ~
T7" a little different between the two, because the operating condition is not exactly the same.
Figure 2 Positionof thermocouples along the pulse tube and the regenerator.
In the
course of the long experimental run, we noticed that the minimum temperature had a change.
The
reason is that this point is situated at the too sharp minimum.
[2] The cold end temperature of the pulse tube changed smoothly with single minimum against the frequency and the orifice valve opening.
But, against the bypass valve opening it showed peculiar variation, which
looked to have a periodicity, one peak for one rotation of the valve handle. a new one with the flow coefficient 34 % larger. rotation of the valve handle was observed.
So, this valve was replaced by
For this valve too, similar periodicity, one peak for one
Next, instead of the bypass valve, we tried to regulate the gas
flow with a series of plates, each of which had a hole with successively larger diameter.
Still, complicated
temperature fluctuation existed, but without clear periodicity in this case. From these experimental facts, we concluded as follows. inside the double-inlet pulse tube.
There should be an unstable circumstance
Small change of the flow coefficient in the bypass valve opening gives a
start for the big variation of the cold end temperature of the pulse tube.
[3] In any way, the anomalous behavior is considered to occur commonly in the double-inlet pulse tube refrigerator inevitably.
Next step is to investigate the influence of the bypass valve opening upon the
temperature distribution along the pulse tube and the regenerator.
So, 6 Au-Fe ( Fe 0.07 at. % ) vs cromel
thermocouples were soldered onto the pulse tube about equal distance apart and 5 thermocouples onto the regenerator as shown in Figure 2.
The typical results are shown in Figures 3 and 4.
The periodic
temperature variation against the handle rotation of the bypass valve is clearly seen for the pulse tube, but only a trace for the regenerator.
The amplitude of the temperature variation is large in the first 4 rotations
around the middle of the pulse tube ( T3 and T4 ).
It could roughly be said that there are 2 kinds of the
temperature distribution, the one with low temperature ( -220 K ) at the position T4 of the pulse tube, which corresponds to the case with the bypass valve closed and the other with high temperature ( -270 K ) at the position T4.
In the transient region, we have large fluctuations of the temperature distribution
against the bypass valve opening.
In some case we obtain a very low temperature ( for example, point A )
ICEC16/ICMC Proceedings (Cv) 40001
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12
25
4
8
12
(Valve handle rotation number) Bypass valve opening Figure 4 Variationof temperature at various positions along the regenerator with opening the bypass valve. Symbols TT,Ts, "'",Tzl correspond to the same symbols in Figure 2. The bottom curve shows the temperature T1 in expanded scale. [Operating condition] frequency: 1.9Hz, orifice valve opening: Cv 0.0297.
in this region, but it is not so stable. Figure 4 shows that the temperatures Ts, T9 and T10 in the regenerator were about 230 K, 270 K and almost room temperature respectively when the cold end temperature of the pulse tube was lowest. a temperature distribution was quite different from our expectation.
Such
The reason for this is not clear at the
moment.
[4] The time variation of temperatures, T~, T2 and T4 along the pulse tube at the optimum condition in the region B is shown in Figure 5 and that at the condition corresponding to the point A in Figure 6. Although all curves in these figures contain similar noise, the curves in Figure 5 show more clear periodicity. This means that the gas flow in the pulse tube is much larger in the case of Figure 5 compared with the case of Figure 6.
266
ICEC16/ICMC Proceedings 184
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Figure 6
Time variation of Tl,T2 and T4.
[Operating condition]
[Operating condition] frequency: 1.925Hz,
frequency: 1.972Hz,
orifice valve opening : Cv 0.0184,
orifice valve opening : Cv 0.0100,
bypass valve opening : Cv 0.0590.
bypass valve opening : Cv 0.0218.
The time variation of temperatures, T8 and T9 in the regenerator is only about 0.3 K.
From this, one
can say that amount of the regenerative materials is sufficient. CONCLUSION An one-stage double-inlet pulse tube refrigerator generated a low temperature 23.9 K. Anomalous temperature variation occurred in the pulse tube against the bypass valve opening, which would be common in this type of the pulse tube refrigerator.
The reason for this is considered as follows.
With opening the
bypass valve, the temperature distribution along the pulse tube changes from the one with low temperature (-~220 K ) at the position T4 of the pulse tube to the other with high temperature ( ~270 K ) at the position T4.
In the transient region do occur large fluctuations of the temperature distribution.
REFERENCES Zhu, S. W., Wu, P. and Chen, Z., Double inlet pulse tube refrigerators 9an important improvement, Cryogenics ( 1990 )30 514-520
Temperature Stability of Pulse Tube Refrigerators
Nobuaki Seki* , Shuichi Yamasaki* , Junpei Yuyama* , Masahiko Kasuya** , Kenji Arasawa ** , Shinji Furuya
and Hidetoshi Morimoto
ULVAC JAPAN, Ltd., 2500 Hagisono, Chigasaki, Kanagawa 253, Japan ** ULVAC CRYOGENICS, INCORPORATED, 1222-1 Yabata, Chigasaki, Kanagawa 253, Japan
This paper describes experimental results comparing four configurations in the two types of pulse tube refrigerators (the double inlet pulse tube refrigerators --DIgrR-and the single orifice pulse tube refrigerators --OPTR). Though temperature of the DIPTR (using needle valves) became unstable within 200 h, that of the OPTR (using either a needle valve or an orifice plate) was maintained stable within +0.5 K. The flow direction dependence of the needle valve impedance in the bypass line is considered to cause a gas flow circulating in the path which is composed of the bypass line, the regenerator and the pulse tube. This is suggested to be the origin of temperature instability in long-term refrigerator operation.
INTRODUC-~ON Temperature instability in double inlet pulse tube refrigerators (denoted by DIgrR) are recently reported by several research groups. We have studied temperature stability of pulse tube refrigerators, and found a temperature instability while a DIPTR was running with 20 W heat load applied to the cold end. On the other hand, the temperature instability has not been observed in a single orifice pulse tube refrigerator (denoted by OPTR) [ 1]. We have examined the difference between the needle valve and the orifice plate in the bypass line. This paper describes experimental results on the temperature instability and discuss its origin.
EXPERIMENTAL APPARATUS The original experimental setup is a DIPTR. A schematic diagram is shown in Figure 1. The refrigerator can be operated in four configurations by slightly modifying the room temperature parts. The four configurations are as follows. 9OPTR using a needle valve in the buffer inlet line. 9OPTR using an orifice plate in the buffer inlet line. 9DIPTR using needle valves in the bypass line and the buffer inlet line. 9DIPTR using orifice plates in the bypass line and the buffer inlet line. The orifices of a diameter less than 1 mm are used in the experiment. Low temperature parts are common for all experiments. Their dimensions are as follows. The inner diameter of pulse tube is 24 mm and the length is 157 mm. The inner diameter of regenerator is 24 mm and the length is 162 mm. Stainless steel screens of 80 mesh and 250 mesh are put in the regenerator. There are heat exchangers at the hot and cold ends of the pulse tube. The buffer volume is about 460 cm 3. A commercial compressor for G-M 267
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ICEC16/ICMC Proceedings
refrigerators is used in this work (C10; ULVAC CRYOGENICS, INCORPORATED). The nominal power consumption of the compressor is 1.5 kW. Charged pressure of Helium gas is 1.72 MPa.
EXPERIMENTAL RESULTS AND DISCUSSIONS Experimental results on the temperature stabilities in four configurations are listed in Table 1. Figures in the last column stand for experimental durations. For example, " stable (> 500 h) " shows that the temperature remained stable over 500 h, and " unstable (< 120 h ) " shows that the temperature became unstable before 120 h passed. This table denotes that the temperature was maintained stable for all configurations except the D W r R using needle valves with 20 W heat load applied to the cold end. The refrigeration power of our OPTR is about 23 W at 100 K. Figure 2 shows the temperature stabilities in two configurations of pulse tube refrigerators with 20 W heat load applied to the cold end. One is the DIPTR using needle valves in the bypass line and the buffer inlet line. The other is the OPTR using a needle valve. Though the temperature of the DIPTR became unstable within 200 h, that of the OPTR was maintained stable within +0.5 K over 500 h. A small spike in the temperature trace was due to transient increase of heat input caused by the failure of the vacuum system. The temperature instability of the DIPTR was repeated in every experimental run when needle valves were used. Figure 3 shows the temperature traces of two configurations of the DIPTR with 20 W heat load applied to the cold end. One is the trace obtained with orifice plates in the buffer inlet line and bypass line. Another is obtained with needle valves, and this trace is shown also in Figure 2. The temperature in the DIgFR using orifice plates was maintained as stable as the OPTR trace shown in Figure 2. Though a needle valve is usually used in the bypass line to control the gas flow rate through it, the needle valve is suspected to cause the temperature instability. Flow impedance measurements were then performed on the valve removed from the refrigerator system. The flow impedance in the normal direction was found to be different from that in the reverse direction. We consider that the flow impedance difference of the valve in the bypass line causes a gas flow circulating in the path which is composed of the bypass line, the regenerator and the pulse tube. Though circulating flow is thought to be much smaller than the main flow generated by pressure oscillation, we think that this is the origin of long-term temperature instability. We have not yet explained why the temperature remains stable over 200 h with a smaller heat load (e.g. 15 W) in the DIPTR using needle valves
CONCLUSION In the DIPTR configuration, the cold end temperature is stable when orifice plates are used, but that is unstable when needle valves are used. The origin of this temperature instability is suggested that the gas flow circulating in the path which is composed of the bypass line, the regenerator and the pulse tube. This circulating flow is caused by the flow direction dependence of the valve impedance in the bypass line. When needle valves are used, the cold end temperature is stable in OPTR, but that is unstable in DIPTR. Thus, the impedance difference of the needle valve in the buffer inlet line does not disturb the gas flow in the refrigerator system.
REFERENCE 1 Yuyama, J., Seki, N., Kanada, T., Kasuya, M., Arasawa, K., Furuya, S. and Morimoto, H. Refrigeration performance and temperature stability in pulse tube refrigerators Extended Abstract of 54th Meeting on Cryogenics and Superconductivity (1995) 68 (in Japanese)
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Compressor
Buffer volume
-500 h) stable (>500 h, >500 h) stable (>200 h) stable (>400 h) stable (>80 h, >200 h) unstable (< 120 h, 400 h) stable (>600 h, >500 h)
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Pulse Tube Refrigerator with Low Temperature Switching Valve
Jingtao Liang*, Cunquan Zhang**, Lie Xu**, Jinghui Cai*, Erchang Luo* and Yuan Zhou* *Cryogenic Laboratory, Chinese Academy of Sciences, PO Box 2711, Beijing 100080, China **Cryogenic Laboratory, Shanghai Jiaotong University, Shanghai 200030, China
A new type of pulse tube refrigerator, termed pulse tube refrigerator with low temperature switching valve, is proposed. It is suitable for industrial applications that require large refrigeration powers. In this kind of pulse tube refrigerator a recuperative heat exchanger instead of a regenerator is used and a switching valve is installed at the cold end of an orifice pulse tube. The adiabatic expansion efficiency of the orifice pulse tube with low temperature switching valve, which actually works as a new type of expander, has been experimentally investigated. Adiabatic efficiencies higher than 40% have been achieved in the preliminary experiments.
INTRODUCTION Pulse tube refrigerator which was invented by Gifford and Longsworth in the 1960's has been developing at an amazing speed in recent years[2-8]. Its refrigeration performance is now becoming comparable to that of G-M refrigerator or Stirling refrigerator. So pulse tube refrigerator is finding more and more applications. However, as a regenerative refrigerator, it has been so far mainly studied and used as a small cryocooler for the applications such as infrared cooling. The cooling powers generally fall in the range of 100 mW to 100 W at liquid nitrogen temperature. In principle pulse tube refrigeration is capable of producing larger cooling powers. Some researchers have studied the possibilities of using pulse tube refrigerator for applications such as small scale natural gas liquefaction. One method often considered is to enlarge the geometric dimensions of a closed-cycle regenerative pulse tube refrigerator to achieve the required refrigeration power. In this paper we propose a new method for achieving large refrigeration powers for some industrial applications. DESCRIPTION OF FLOW DIAGRAM The flow diagram of the conventional regenerative pulse tube refrigerator may be inadequate for the applications which require large refrigeration power for the following reasons: 1. It is not suitable for working in the two-phase region of the working gas. In the applications such as gas separation and liquefaction, especially in the case that there is a source of pressurized gas, it is often convenient to use the gas to be processed as the working gas in pulse tube. Since the high pressure gas flow and the low pressure gas flow alternatively pass through the same channel in the regenerator, it is not easy to separate liquid produced. Liquid in the regenerator may increase the flow resistance and reduce the expansion efficiency of the incoming gas flow. 2. The void volume of regenerator which has to be pressurized and depressurized is quite large. It greatly reduces the pressure ratio in pulse tube. 271
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Therefore, based on our previous work[6-8] we propose to replace the regenerator with a recuperative heat exchanger and to install a low temperature switching valve at the cold end of the orifice pulse tube. Such a refrigeration system, termed pulse tube refrigerator with low temperature switching valve, is schematically shown in Figure 1. The switching valve alternatively switches the cold end of LP
HP
__.Gas reservoir ~
Recuperative heat exchanger
~
Orifice Hot end heat exchanger Flow straightener
'
Pulse tube
Flow straightener Gas-liquid separator
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1 Figure 1 Flow diagram of pulse tube refrigerator with low temperature switching valve pulse tube to the high pressure channel and the low pressure channel. The incoming high pressure gas stream is cooled down by the low pressure gas stream in the recuperative heat exchanger. Then it goes into the pulse tube, via a gas-liquid separator to eliminate liquid drops if necessary, expands to lower pressure and leaves pulse tube with lower temperature, thus produces cooling. If part of the gas is liquefied after expansion in pulse tube, the low pressure gas stream passes through a gas-liquid separator. The liquid part is removed and the exhaust gas stream is warmed up in the recuperative heat exchanger. Finally it goes back to the compressor or gets out of the system. This kind of pulse tube refrigerator can be readily industrialized due to the following reasons: (1) the recuperative heat exchanger and valved compressor are commercially available; (2) the low temperature switching valve which is almost the same as that used in a piston expander is a well developed industrial technology; (3) pulse tube is facile to fabricate and is flexible for spatial arrangement. In fact the orifice pulse tube coupled with switching valve works as a new kind of expander except that heat instead of work is extracted. It has the advantage of involving no reciprocating piston and no high speed rotating components at low temperature, although it has a low temperature switching valve. The adiabatic efficiency of the orifice pulse tube expander is the crucial point for its application prospect. Experiments have been conducted to investigate the adiabatic efficiency. EXPERIMENTS If the gas enters the pulse tube at the pressure P~ and the temperature T~, and after expansion its pressure and temperature become P2 and T2, the adiabatic efficiency can be expressed as
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rl s= (T,- T2) / (T,- T2' ) = A T / A T'
(1)
where T2' is the temperature that can be achieved when the gas is isoentropically expanded from state (P,, T,) to P2. It can be known from the properties of the working gas. So to know rl~ we need to measure P,, T,, P2, and T 2. A test bench has been set up. A membrane compressor of 22 Nm3/h is employed to supply the pressurized gas. Two large buffers are installed at the inlet and outlet of compressor. The working gas is pure air. For the preliminary experiments, a low temperature switching valve dismantled from an old helium liquefier is used. Three pulse tubes have been tested. They are 400 mm long and of 14 mm, 19 mm, and 24 mm ID respectively. The heat exchanger at the hot end of pulse tube contains about 60 stainless steel tubes of 2 mm ID and 200 mm in length, which are cooled by water. The orifice is an adjustable needle valve. The compressed air enters the pulse tube at ambient temperature, expands to 1 bar and directly goes back to the buffer volume connected with the inlet tube of compressor. No heat exchanger is used to precool the inlet gas flow for the pulse tube and to recuperate the cold of the outlet gas flow in the preliminary experiments. The purpose of the experiments is to study the adiabatic expansion efficiency. The cold parts are thermally insulated only with perforated plastics. Pressures in pulse tube, gas reservoir, and the inlet and outlet tubes are measured with pressure sensors and recorded with a computer. The inlet and outlet temperatures are measured with copper-constantan thermocouples. In the experiments we have measured the adiabatic efficiency under various conditions. The influences of pressure ratio, frequency, orifice opening, and pulse tube diameter on the adiabatic efficiency are investigated. As an example of the experimental results obtained, Figure 2 shows the 45 f = 2.6 Hz 40
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P,/P2 Figure 2 Adiabatic expansion efficiency vs pressure ratio adiabatic efficiency in function of the pressure ratio at different orifice openings. The frequency is 2.6 Hz It can be seen that with proper opening of the orifice the adiabatic efficiency generally increases with increasing pressure ratio. It can also be seen that the orifice and gas reservoir can effectively increase the adiabatic efficiency. At the orifice opening of three turns and the pressure ratio of 11 bar/1 bar, an adiabatic efficiency of 42% is achieved, the corresponding temperature drop being 63.6 K.
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CONCLUSIONS On the basis of the requirements of certain industrial applications, a new type of pulse tube refrigerator, called pulse tube refrigerator with low temperature switching valve, is originally introduced. The concept of orifice pulse tube expander is proposed. Preliminary experiments have been carried out to investigate the adiabatic expansion efficiency of orifice pulse tube. Experimental results well demonstrate the feasibility and the application potential of this new type of pulse tube refrigerator. ACKNOWLEDGMENT This work is supported by the National Natural Science Foundation of China under No.59506003. REFERENCES 1 Gifford, W.E. and Longsworth, R.C., Pulse tube refrigeration, Trans. ASME, J. Eng. Ind. (1964) 86 264 2 Mikulin, E. I., Tarasov, A. A., and Shkrebyonock, M. P., Low temperature expansion pulse tubes, Adv. Cryo. Eng. (1984) 29 629 3 Liang, J., Zhou, Y., and Zhu, W., Development of a single-stage orifice pulse tube refrigerator capableof reaching 49 K, Cryogenics (1990) 30 49 4 Zhu, S., Wu, P., and Chen, Z., Double inlet pulse tube refrigerators: an important improvement, Cryogenics (1990) 30 514 5 Cai, J.H., Wang, J.J., Zhu, W.X., and Zhou, Y., Experimental analysis of multi-bypass principle in pulse tube refrigerators, Cryogenics (1994) 34 713 6 Liang, J., Ravex, A., and Rolland, P., Study on pulse tube refrigeration. Part 1: thermodynamic nonsymmetry effect, Cryogenics (1996) 36 87 7 Liang, J., Ravex, A., and Rolland, P., Study on pulse tube refrigeration. Part 2: theoretical modelling, Cryogenics (1996) 36 95 8 Liang, J., Ravex, A., and Rolland, P., Study on pulse tube refrigeration. Part 3" experimental verification, Cryogenics (1996) 36 101
On-off Timing Computer Control System for Valved Refrigerator
Li Zhao, Tao Sun, Jianyao Zheng, Zhixiu Huang, Guobang Chen Cryogenics Laboratory, Zhejiang University, Hangzhou 3 10027, P.R.China
The essay focuses on the on-off timing computer control system with which the operating frequency and on-off time ratio can be adjusted freely. In four-valved refrigerator control system, phase difference of pressure waves can also be changed easily as the corresponding control program slightly touched. It brings possibility to carry out the phase shift mechanism of pulse tube refrigeration, regulate the operating parameters close to the optimum value, and eventually obtain the similar performance of G-M refrigerator.
INTRODUCTION The generic family of cryocoolers which employ valves to control the gas supply system such as G-M, Solvay, Postle engine and so-called valved pulse tube refrigerator should experience a process of sequence optimization for best performance. The pulse tube refrigerator which is eliminated the moving displacer in cryogenic temperature range has the potential for lower cost, higher reliability and less vibration at the cold tip[ 1]. Its advantages over other cycles arouse tremendous interest in fabricating progressed prototypes and in demonstrating no-load temperature drop. Still, the refrigerator's prospects for spacecraft and commercial application are restrained due to the limitation of cooling capacity. Therefore, valved pulse tube refrigerator emerges and by proper relative timing is latent to provide greater refrigeration power, even obtains the comparable performance with that of G-M refrigerator[2]. Consequently, it is likely a digital control system will receive development attention for fluidic control of the inlet/outlet valves.
SYSTEM REPRESENTATION Emphasis has been placed on the digital control system for correct phasing to achieve the desired refrigeration cycle. The on-off timing computer control system we designed is an open-loop control system which is calibrated to yield needed results[3]. In other words, when a hot-key touch changes the operating frequency, phase angle or on-off time ratio, a mark is then made on the screen corresponding to this alteration as well as the real time display of the solenoids by four light emission diodes. Then, the data are transmitted simultaneously from a parallel port and fed into the circuit to energize the solenoid driver by applying rated voltage to the coil, which is used to provide timing of solenoid in control circuit. The process is continued until all required values for sequence are determined. Figure 1 illustrates the functional block diagram of the system. The circuit is divided into low voltage part which is connected to the computer by a conversion cable and high voltage one which is wrapped by insulation layers. The isolated approach is an advantage when risk of shock, fire or disturbance must be minimized. 275
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SOFTWARE DESIGN Scheduling timing is handled by an interrupt-driven program whose control-oriented flowchart is graphed in Fig.2. At the beginning of the program, save the state of the current process in its descriptor. That is, the status of the processor at that time is saved on a push-down stack at which the machine-register stack pointer is pointing. Then, the interrupt occurs with initialization of the screen and the 8253 programmable interval timer. Because all required responses are time critical, the timing task execution is imer-linked with the tick, an elapsed time counter updated by interrupts from the system real-time clock implemented with the 8253[4]. At that moment, the tick's handler is loaded up and set going after all information is saved. Soon afterward, the system gets hot-key-touch instructions since the operator needs to have access to all system information via keyboard/display unit, until it is ordered to stop running. We then 'plant' the necessary reference to the program location that the compiler fixes at an absolutely defined address in the interrupt vector area to end up. We employ C++ for the software design because of its object-oriented program technique, language flexibility, program readability and its simple interfaces to assembly language routines for execution speed[5]. The range of the operating frequency and on-off ratio can be varied to meet the requirement of distinct pulse tube refrigerator, G-M or Solvay refrigerators by upgrading and modifying the program.
COMPARISON TO TIMER CONTROL SYSTEM By contrast, the on-off timing computer control system excels the timer control system on the following aspects[2]: 1.It exhibits real-time control with fast response. Its minimum time delay is 5ms less than that of the timer control system. 2.It offers the user a friendly interface to an otherwise complex subject, on-off timing, with the lighting emission diodes to predict system performance and with the screen to indicate the time sequence, while the timer control system needs a round-blade screw-driver or an Allen wrench to turn the operation mode or to change time unit and rated time as well as calculation by heart. 3.In today's marketplace connectivity is perhaps the obstacle to product success. The system is more practical for it can switch from controlling two-valved pulse tube to four-valved one by simply modifying several lines of instructions in the program while the hardware remains unchanged. This principle is finding enormous use today in other types of valved refrigerators such as G-M and Solvay refrigerators. 4.It claims low cost and great potential for commercial application. Though the minimum amount of the OMRON mode H3BA solid-state timers used in the timer control system is three, it can't extend the open period of the second valve beyond half cycle and can't adjust the interval time of the two valves either. Therefore, the system utilizes four timers to properly regulate two solenoid valves and seven to control four, which increases the expense terrifically.
APPLICATION Although the G-M cryocooler has passed through the pioneering stage to become well established, its intake/exhaust valve sequence still needs to be optimized, since its cooling capacity under different refrigeration temperatures is to a certain degree determined by the corresponding open angle and interval[6]. However, altering rotary valves is rather troublesome and time-consuming. Hence, instead of testing on piles of rotary valves with different versions, the on-off timing computer control system provides an easy-to-use interface for adjusting the operation parameters and saves great effort concentrated principally on the manufacture of rotary valve varieties. Once the perfect sequence is obtained, the
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corresponding rotary valve can be consequently put into production. Therefore the system exhibits its advantages to bridge the gap between laboratory work and commercial application. The case in point is Cryogenics Laboratory, CAS who employs the system in which the regulation scope of each valve is quite extensive so that the research on optimization of four-valved G-M cryocooler can be readily carried out. Similar to G-M engine, in valved pulse tube refrigerator, proper phasing of the solenoids is also significant but with distinct time sequence. The interval between intake/exhaust stroke is arbitrated to be large than 10ms in order to avoid the mixed up of the compressed/expanded fluid, which commands the system response speed to meet the requirement. Since the delay of its slowest component -- the solenoid driver is less than 5ms, the real time project is designed to satisfy this time criterion. We adopt the device in our three-stage pulse tube refrigerator whose typical time sequence is shown in Fig.3 and find that impact of the sequence varies according to orifice and by-pass valve open ratio, configuration and strongcoupled operating parameters. For example, extending the exhaust stroke contributes to system performance, especially to the cooling rate of the pulse tube at first and second stage. On the sequence varied point in Fig.4 where the exhaust stroke extends from 0.17s to 0.44s the cooling rate has an abrupt drop.
CONCLUSION The on-off timing computer control system which is able to sequence the solenoids for corresponding motions concerns with the cryocoolers having valves to control the flow of working fluid and accelerates the optimization process of their time sequences.
ACKNOWLEDGMENT The project is sponsored by the National Natural Science Foundation of China.
REFERENCES Radebaugh, R., A review of pulse tube refrigeration, Adv. Cryo. Eng., Vol.35, (1990) Chen, Guobang et al, Comparison test of two-valved and valveless pulse tube refrigerator, Cryogenics, (1994) Vol 34 ICEC Supplement, 151-153 Pericles Emanuel and Edward Left, Feedback control system, McGraw-Hill Book Company, New York(1990) Intel component data catalog, Intel Corporation, California(1982) Cooling, J.E., Software design for real-time system, Chapman & Hall World Publishing Corp, UK(1991) Li,R. and Onishi,A. Optimization of intake/exhaust valve for 4K Gifford- McMahon cryocooler, Proceedings of ICEC- 16, Japan(1996), to be published
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INTERRUPT SCREENFORMAT[ SOLENOIDS
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; Pin the pulse tube (e.g. [1, 3]). It has been argued that < ISI> in the 4 K regime is greatly reduced due to the fact that the amplitude Ta of the oscillating fluid temperature is strongly diminished in the non-ideal 4He gas [2, 5]. However as shown below, the pressure dependence of H(p, T) in the real gas results in < 111> > 0 even when the dynamic temperature amplitude Ta drops to zero. From general thermodynamic relations the change in specific enthalpy dh(p, T) of the working fluid can be written as dh(T, p) = Cp d T + pl(1 - [3T) dp, where Cp, P, and 13 are the coefficient, respectively. For amplitudes Ta and pa = Ap/2. to < I~I > = ~ Td ( r n by
(J)
specific heat at constant pressure, the density, and the volume expansion small pressure variations dT and dp can be approximated by the dynamic For an ideal gas 13 - 1/T and then the 2nd term in Eq. (1) is zero, which leads mass flow rate). For adiabtic pressure variations in the pulse tube Td is given
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Td "" (13T/(pCp)) Pal.
(2)
Calculations Of Td at the pulse tube cold end for
= 19.5 bar and Ap = 9.6 bar by use of the NIST12 4He-database show that Td decreases rapidly below 10 K, which is due to the anomalously large volumetric specific heat 9Cp, and it even drops to zero for T = 2.0 K, since 13 disappears. The adiabatic enthalpy variation Ah = dh, however, still remains. This is seen by inserting Eq. (2) into (1) leading to the simple relation Ah = 9 1 Pa 9
(3)
It follows from Eq. (3) that < H > - 1/9 = dh/dp. For 4He below 4 K the density and thus the average enthalpy flow < I2I > is only weakly T-dependent. Figure 4 illustrates the variation of CpTd (first term in Eq. (1)) and Ah (Eq. (3)) below 20 K. Clearly, below 10 K the adiabatic enthalpy variation in the real 4He fluid is higher than in the ideal gas. 12
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= 19.5 bar, Pd = Ap/2 = 4.8 bar. In conclusion, the present work demonstrates that a multi-stage PTR working with 4He gas can reach temperatures near the X-line. This result is of fundamental interest concerning the limits of pulse tube refrigeration. Future developments will aim towards higher net cooling powers around 5 K, as required for most practical applications.
ACKNOWLEDGEMENTS We thank U. H~.fner (Leybold) for providing the magnetic regenerator material. This work is supported by the German Ministry of Science and Technology (BMBF) under contract no. 13N6513.
REFERENCES 1 Matsubara, Y. and Gao, J.L., Novel configuration of three-stage pulse tube refrigerator for temperatures below 4 K Cryogenics (1994) 34 259 2 Matsubara, Y. and Gao, J.L., Multi-stage pulse tube refrigerator for temperatures below 4 K, In: Cryocoolers 8 Plenum Press, New York (1995) 345 3 Zhu, S., Wu, P., Double inlet pulse tube refrigerators: an important improvement Cryogenics (1990) 30 514 4 e.g. Kuriyama, T. et al. High efficient two-stage GM refrigerator with magnetic material in the liquid helium temperature region Adv. Cryog. Eng. (1990) 35B 1261 5 Thummes, G., Bender, S. and Heiden, C., Approaching the 4He lambda-line with a liquid nitrogen precooled two-stage pulse tube refrigerator Cryogenics (1996), to appear
Thermodynamic Calculation of Three-stage Pulse Tube Refrigerator*
Guobang Chen, Zhongming Xia, Li Zhao, Liming Qiu, Jiangyao Zheng, Jianping Yu Cryogenics Laboratory, Zhejiang University, Hangzhou 310027, P.R.China
The refrigeration capacity and thermodynamic losses of a three-stage pulse tube refrigerator have been calculated by means of the separate method of gas piston and its boundary layer. The influence of configuration and operating parameters on refrigeration performance is discussed.
INTRODUCTION Past ten years saw fantastic spurt in development of pulse tube refrigerator. The case in point is the liquid helium temperature reached in multi-stage pulse tube refrigerator[ 1]. Otherwise, its theoretical investigation isn't satisfactory, especially in multi-stage pulse tube. In this paper, the method which is always adopted in calculation of cooling capacity in G-M refrigerator is applied to count the theoretical cooling capacity of valved three-stage pulse tube refrigerator. Furthermore, main thermodynamic losses in the procedure are discussed. Finally, the net cooling capacities at each stage are obtained.
THERMODYNAMIC CALCULATION METHOD Thermodynamic calculation methods of a cryocooler include simultaneous and isolated ones. In isolated method, the premises are that irreversible losses do not exert any influence on ideal cycle and do not have any connection with each other, so they can be counted independently. Although the former which takes cycle and all kinds of losses into consideration is more accurate than the latter, it hasn't found wide application in engineering due to its tremendous work load. The application of isolated method in three-stage pulse tube refrigerator is consist of three steps: firstly, refer to certain physical model (isothermal or adiabatic model) and count the theoretical cooling capacity Q~; then, calculate the losses in the refrigerator respectively; finally, the actual cooling power Q~ is obtained.
(1)
Q~j - Q~j - ~ AQj + ~-] AQj+~ j
j+l
in which, Q~j represents the actual cooling power in pulse tube at j stage, Q~j is the theoretical cooling power at j stage, ~ AQj and ~ AQj+ 1 a r e the sums of irreversible losses at j stage and transferable irreversible J"
j+l
losses at j+l stage as its gain cooling. In order to carry out thermodynamic calculation, gas piston and trapezoid pressure wave in the pulse tube are basically hypothesized. The gas piston has the similar function to the solid displacer in G-M refrigerator. * The project is supported by the National Natural Science Foundation of China. 315
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THEORETICAL COOLING CAPACITY The shape of equivalent P-V diagram of the cold volume in a valved pulse tube refrigerator is similar to that of a G-M refrigerator, so its total area can be counted by means of calculation used for G-M refrigerator[2]. Under steady condition, the system energy equilibrium equation to the cold ends of pulse tube and egenerator is, (2)
Qc - ~dW
while the net work which is done by the cold volume to exterior ~ d W should be equal to the area of the equivalent P-V diagram. Therefore, the theoretical cooling capacity Q~ in a valved pulse tube can imitate the calculation of that in G-M refrigerator: Q~ - ~ d W - f (Ph - Pt)V~
(3)
where, V~ is the maximum of cold volume. The cold volume in pulse tube can be considered as the maximum volume of gas flowing into the pulse tube at the compression process.
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(A) (B) (C) (D)
The construction of regenerator EraNi only Ero. ~Ybo. INi(50%)+EraCo(50%) Er(Coo. 2Nio. s)(25%)+Ero..~Ybo. INi (25%) +Era Co (50%) Ero. 9Ybo ,Ni (50%) +Era Co (25%) + Ero. 7sGdo. zsNi(25%)..
Table III The refngeration capacities of the GM refngerators with above four types of regenerators at 4.2K as a function of reciprocating speed. reciprocating refrig.power refrig.power refrig.power refrig.power
speed of type(A) of type(B) of type(C) of type(D)
24 0.88 0.96 1.0
. .30 . . . . 36 42 0.66 ( w a t t . ) 0.96 1.05 1.03 1.10 I. 12 1.17 1.17
REFERENCE 1.
2. 3. 4. 5. 6.
For instance, Hashimoto, T., Li ,R., Matsumoto, K., Sahashi, M., Y ayama H., and Tomokiyo, A., Recent Progress in the Materials for Regenerator in the Range from 4.2K to 20K, Prec. Intern. Cryogenic Material Conf. p.667, 1988, Shengyan China. Nagao, M., Inaguchi, T., and Yoshimura, H., Helium Liquefaction by Gifford McMahon Cycle Cooler, Advances in Cryogenic Engineering 35(1990) 1183. Kuriyama, T., Hakamada,R., Nakagome, H., Tokai, Y., Sahashi, M., Li, R., Yoshida, O., Matsumoto, K., and Hashimoto, T., High Efficient Two-stage GM Refrigerator with Magnetic Material in the Liquid Helium Temperature Range, Advances in CD'ogenic Engineenng ( Plenum Press, New York ) 35(1990) 1261. Debye, P., Ann.Physik 39(I912)789. Hashimoto, T., et al.,Teion-Kougaku 29(1994)51. Seshake, H., et al.,Advances Cryogenic Engineering 37(1992)995. Yabuki, M., et al., Recent Progress on Application of High Entropy magnetic matenals to the regenerator in He Temperature Range, Proc.7th Intern.Cryocooler, p.605,Kartland AIZB NM,(1993). Tsukagoshi, T., et al., to be published to Cryogenics in 1996.
Development of a 4K GM/JT Refrigerator for Maglev Vehicle
Satoru Fujimoto, Shoichi Taneya, Toshiyuki Kurihara, Katsuya Miura, Keiji Tomioka Masakazu Okamoto, Tatsuya Yamaguchi, Shinichi Kasahara Motoaki Terai*, Akihiko Miura*, Hiroyuki Nakao**, Takahiro. Fujinami**, Yasutoshi Nakamoto*** Daikin Industries Ltd, 1304 Kanaoka-cho,Sakai,Osaka,Japan *Central Japan Railway Company, 1-6-6 Yaesu,Chuo-ku,Tokyo 185,Japan **Toshiba Corporation, 1 Toshiba-cho,Fuchu,Tokyo 185,Japan ***Toshiba Corporation, 1-1-1 Shibaura,Minato-ku,Tokyo 105-01,Japan
INTRODUCTION In our project, a Gifford-McMahon/Joule-Thomson(GM/JT) refrigerator for Maglev vehicle has been developed in partnership with Central Japan Railway Company and Toshiba Corporation. Commercial GM/JT refrigerators had been used for radio astronomy and other laboratory systems, however they had not been applied to Maglev vehicle because of the small refrigeration capacity due to the low efficiency. On-board helium refrigerators require large capacity, high efficiency and reliability even under severe environmental conditions. Major efforts have been made to improve the refrigeration capacity and efficiency with considering reliabilities.
SPECIFICATIONS Principal Specifications of an on-board refrigerator are listed in Tablel. The refrigeration efficiency had to be increased approximately 80%. In addition, the refrigerator for Maglev vehicle has to be operated under very restricted environment. The constraints for designing the refrigerator are listed below. (1) The refrigerator has to be stored in a limited space and it can't exceed the certain weight limit (2) The power consumption of the refrigerator is restricted. (3) The refrigerator has to withstand the vehicle vibration, magnetic field and variable ambient temperature. (4) The refrigerator has to deal with the excessive amount of heat load momentarily when the superconducting magnet is energized or de-energized Table l Principal specifications of helium refrigerator Refrigeration capacity Input power COP Vibration condition Magnetic field condition Weight
for Maglev 8W (at 4.4K) 8kW 1.0X10-3 exp:-4-10G, comp :4- 5G max 0.1Tesla < 300kg
commercial (CG308) 3.5W (at 4.3K) 6.4kW 5.5X10-4
255kg
STRUCTURE OF REFRIGERATOR The refrigerator consists of an expander unit, a compressor unit and a flexible tube that connects these units. Fig.1 Shows the flow diagram of refrigerator. 331
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The structure of the expander unit, which is installed on a side of the helium tank on top of the superconducting magnet, is shown in Fig.2. It has a 2-stage GM expander(precooler), three JTheat exchangers, a JT valve, and a radiation shield which will be cooled by 1st stage of G M expander. The displacer inside of expander is driven pneumatically at 2Hz by switching high/low pressure with rotary valve. The major components of the compressor unit equipped on bogie frame has two scroll compressors, oil separators, an adosorber, and an air-cooler for cooling helium gas/oil. They are fixed inside of the aluminum frame like Fig.3. These compressors have different scroll forms which were improved from what have been used for air conditioning. From this fact, it is obvious that the compressors hold high reliability and cost performance.
DEVELOPMENT OF REFRIGERATOR There were mainly three aspects to improve refrigeration capacity and efficiency. (1) Optimizing the pressure condition in refrigeration system (2) Improvement of compressor (3) Improvement of GM expander
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(1) Optimizing the pressure condition It is important to match the compressor's property with the expander's. Because of the fact that the camot efficiency of GM expander changes with pressure since refrigeration occurs in principle of Simon expansion, not adiabatically, it is necessary to design the refrigerator by optimizing the pressure condition. Fig.4 shows the input consumption with constant efficiency of compressors and expander. The pressure condition is determined by intermediate pressure because high and low pressure are fixed by the control valves. When the intermediate pressure is high, the pressure ratio of second compressor becomes small and the input consumption of second compressor will decrease. In the system used, the intermediate pressure is set at 8-9kgf/cmZG to optimize the properties of compressors and expander. In order to adjust this pressure, the expansion volume had to be extended, and suck-in volume of the 2nd compressor had to be enlarged. (2) Improvement of compressor Scroll compressor used for the project is called "fixed compression ratio" type, and its best compression ratio depends on the number of scrolls. If the actual ratio changes, the compression efficiency will decrease because of over-compression or re-compression. The scroll are optimized to get best efficiency as described in the section of "Optimizing the pressure condition" above. As a result, the compression efficiency increased approximately 10% on first compression and 15% on second compression. (3) Improvement of GM expander GM expanders with new regenerator using rare-earth materials such as Er and Ho, have been developing and their capacity and efficiency have shown tremendous improvement in recent years. The specific heat of lead which is a conventional regenerator becomes smaller when temperature is lower than around 15K. On the other hand, new regenerators have an outstanding peak of specific heat between 10K to 20K that is mainly caused by magnetic phase transition. Therefore, Er3Co compound is adopted for the second stage regenerator of GM expander. Er3Co has high specific heat property within the range of 2nd stage temperature. The quality control when they are on manufacturing process is relatively high, and finished product has rigidity. In the range of second stage temperature, 55% of the capacity shown on the P-V diagram is lost and the rest is remained the refrigeration capacity. Some of these losses are caused by shuttle, pumping and so forth, however the largest loss comes from the inefficiency of regenerator. It is estimated that 20% of the regenerator loss can be regained by utilizing Er3Co (Fig.5). Furthermore, the computed simulator that can calculate P-V diagram and some losses of GM expander has been developed. By using simulator, some parameters to design the refrigerator are determined such as displacer stroke, regenerator quantity, and the compound size.
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CHARACTERISTICS OF REFRIGERATOR (1) Frequency-Capacity property The refrigeration capacity is 8.3W when the input consumption is 7.9kW, and the COP is 1.05x10 -3 at 53Hz. The temperature in helium tank is below 4.5K, and the JT mass flow is 12.512.7Nm3/h. However the capacity without vibration-proof shows 8.5-9.2 watt and the COP i s 1.1x10 -3 at 50-55Hz. This difference of capacity is conductive loss caused by supporting equipment which fixed some elements. COP will decrease at over 55Hz because the capacity of precooler doesn't increase in proportion to JT mass flow. (2) Gas Withdrawal and Load-changing property n| The refrigerators withdraw the helium gas at t 60W/2min" 60W/2mi 70Hz to control the helium tank pressure when ~ H~II Ile~ld superconducting magnets are energized or de, i ! 6W . . . . o!| e~.ergized~ AI].d the refrigerators restrain the i~Iput mode ' ~~.~.~ ~-i ~ ~ ~~.~" consumption and protect the pressure drop in ' [~()~ual(53Hzi , Ehergize helium tank at 40Hz when the heat load is lower En~gize(70Nz) [--]Restraint(z~(JHz) D ~ -i(,TOHz-) or Maglev vehicle stops for more than a certain ._. period of time. The Refrigerators are ordinally ~ 0.:3 I r----g operated in "usual mode" at 53Hz, and can be =~ 0.2 switched from "usual mode" to "restraint mode ~ 0.1 by detecting buffer tank pressure. Fig.6 shows the result of cycle mode test that .~ 15 has two energizing/de-energizing operations each E 10 in a day and switching automatically between "usual mode" and "restraint mode". Pressure in ',,' 5 helium tank is controlled below 0.39kgf/cm2G at 0 0 6 12 18 energizing/de-energizing and 0.2-0.3kgf/cmZG at Time(hour) the other modes.
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has the refrigeration capacity rating under vibration, magnetic field and refrigeration systems are performed to improve the reliability and the
ACKNOWLEDGEMENT This work is supported by Central Japan Railway Company and Toshiba Corporation.
REFERENCES 1) Fujimoto S., Taneya S. and Tamtani I., Development of a GM refrigerator for Maglev(1) 47th Meeting on Cryogenics and Superconductivity (1992) 2) Fujimoto S., Taneya S. and Kurihara T., Development of a GM refrigerator tbr Maglev(2) 49th Meeting on Cryogenics and Superconductivity (1993) 3) Terai M., Inadama S., Tsuchishima H., Suzuki E. and Okai T., Development of new superconducting magnets for the Yamanashi Test Line Maglev'95 14th International Conference on Magnetically Levita.te~ Systems Bremen'Germany (1995) 267-273
Development of 2W Class 4K Gifford-McMahon Cycle Cryocooler
Inaguchi Takashi, Nagao Masashi, Naka Kouki, Yoshimura Hideto Advanced Technology R&D Center, Mitsubishi Electric Corporation, Tsukaguchi-Honmachi 8-chome, Amagasaki, Hyogo,661 Japan
This paper describes the principal design features and performance of the GiffordMcMahon cycle cryocooler by which we could obtain the cooling capacity of 2.2W at 4.2K. The main features of this machine are its large size expansion space, the use of a regenerator packed rectifiable meshes, and the combinational use of Er3Ni and ErNio.gCo0a as regenerator materials.
INTRODUCTION Since the success of helium liquefaction by the Gifford-McMahon cycle cryocoolcr(hereinafter called 4K GM cryocoolcr)in 1989[1], the 4K GM cryocoolcr has been given attention as a cryocooler which will replace the JT cryocooler because the 4K GM cryocoolcr has high reliability and very easy handling[2]. In particular, 4K GM cryocoolcrs actively have been applied to superconducting magnets such as helium free superconducting magnets[3,4,5] and MRI magnets[6] which don't have to supply liquid helium. However, the use of the 4K GM cryocooler has been limited to small magnets or magnets whose heat load is small. In order to further extend the range of applications of the 4K GM cryocooler, the cooling capacity of the 4K GM cryocoolcr must be improved. During this experimental period we obtained a cooling capacity of 2.2W at 4.2K. The main features allowing us to achieve this result are the machine's large size expansion space, the use of a regenerator packed rectifiable meshes, and the combinational use of Er3Ni[7] and ErNio.gCo0.~ [8]as regenerator materials. In this paper we report on the experimental apparatus used and the results. EXPERIMENTAL APPARATUS Figure 1 shows a schematic of the expander of a two-stage 4K GM cryocoolcr which is an experimental machine. The 4K GM cryocooler has two cooling stages through which heat loads are transferred into the cylinder. Displacers which contain regenerators reciprocate in the cylinder. By the displacers and the cylinder two expansion spaces and a room temperature space are formed. The first regenerator has a twolayer structure. We stacked phosphor bronze screens in the high temperature part and lead shots in the low temperature part. In the second regenerator we packed Er3Ni, or the combination of Er3Ni and ErNio.9Co0.1 as regenerator material, then the effect on cooling capacity was investigated. Figure 2 shows a schematic of the experimental apparatus. The cooling stage of the 4K GM cryocooler was installed in a vacuum chamber. A radiation shield was attached to the first stage, and it enclosed the second stage to prevent the radiant heat from room temperature from entering the second stage. The temperature of the first stage was measured using a Pt-Co resistance sensor and the temperature of the second stage was measured using a carbon glass resistance sensor. A cartridge heater was installed at each stage to measure cooling capacity. In order to secure enough flux to the expander, two compressors (each with rated input power of 6kW) were arranged in a row. For this experiment we kept the differential pressure of the room temperature space at 1.2MPa"-1.3MPa by adjusting a bypass valve which was installed between the high pressure pipe and the low pressure pipe. A pressure transducer was installed into the room temperature space of the expander and a linear 335
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Figure I Schematic of expander of two-stage 4K GM cryocooler.
displacement converter was set up on the drive mechanism of the displacers in order to measure the indicated work. Pressure transducers were also installed at the gas exit and entrance of the expander, and inflowing or outflowing gas pressure was measured. EXPERIMENTAL RESULTS Effect of Volume of Expansion Space on Cooling Capacity Table 1 shows the main parameters of the 4K GM cryocoolers which were used in this experiment. We changed the diameters of the second expansion space from 25.4mm to 60mm in order to investigate the effect of volume of expansion space on cooling capacity. The diameter of the first expansion space of #1 through #3 were 61.9mm; that of #4 was 70mm; and that of #5 was 90mm. The stroke of the cryocoolers were all 32mm. We employed Er3Ni as the regenerator material. The cycle frequency employed was the cycle frequency at which cooling capacity at 4.2K became optimum. The optimum cycle frequency was 45rpm in the cases of # 1 ~ # 3 and the optimum cycle frequency was 33rpm in the cases of # 4 " # 5 . Figure 3 shows the effect of volume of the expansion space on cooling capacity at 4.2K. When the volume of the expansion space is 16.1cm 2 (diameter: 25.4mm), the cooling capacity is 0.34W. The cooling capacities increase in accordance with the increase of volume of the expansion space. When the volume of the expansion space is 90.5cm 2 (diameter: 60mm), the cooling capacity becomes 1.4W. The optimum cycle frequency decreases from 45rpm to 33rpm in accordance with the increase of the volume of the expansion space. The rate of increase of cooling capacity is not therefore proportional to the rate of increase of volume.
Table 1 Main parameters of 4K GM cryocooler Diameter of Diameter of Stroke Optimum cycle 2nd expansion 1st expansion (mm) frequency(rpm) space (mm) space (ram) #1 #2
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Figure 4 (a) Schematic of conventional regenerator and (b) schematic of regenerator packed rectifiable meshes in two additional places of regenerator at .equal intervals Effect of Regenerator Packed Rectifiable Meshes on Cooling Capacity We used two kinds of regenerators shown in Figure 4(a) and 4(b) and investigated the effect of cycle frequency on no load temperature. Figure 4(a) shows a conventional regenerator. Rectifiable meshes are packed only at the ends of the regenerator. Figure 4(b) shows a regenerator packed rectifiable meshes not only at the ends of the regenerator, but also in two additional places of the regenerator at equal intervals. The rectifiable meshes were put between two felt mats to prevent regenerator material from blocking meshes. In both cases Er3Ni was employed as the regenerator material. The experimental machine employed was the 4K GM cryocooler shown in #5 of Table 1. Figure 5 shows the effect of cycle frequency on no load temperature. When we employed the regenerator in Figure 4(a), the no load temperature was 2.78K at the cycle frequency of 30 rpm and it was 4.47 K at the cycle frequency of 45 rpm. The temperature rise was 1.68K. We considered the cause of temperature rise would be that helium gas in the expander flows on one side of the regenerator in the case of the large size 4K GM cryocooler like #5 of Table 1. When we employed the regenerator in Figure 4(b), the no load temperature was 2.63K at the cycle frequency of 30rpm and it was 2.81K at the cycle frequency of 45 rpm. The temperature rise was only 0.18K and it is considered that the rectifiable meshes in two additional places of the regenerator could prevent helium gas in the expander from flowing on one side of the regenerator. Figure 6 shows the cooling capacity of the 4K GM cryocooler using the regenerators shown in Figure 4(a) and 4(b). When we employed the regenerator shown in Figure 4(a), the optimum frequency was 33rpm and the cooling capacity at 4.2K was 1.4W. When we employed the regenerator shown in Figure 4(b), the optimum frequency was 42rpm and the cooling capacity at 4.2K was 1.8W. The optimum cycle frequency of the 4K GM cryocooler using the regenerator shown in Figure 4(b) was 1.3 times higher than that of the 4K GM cryocooler using the regenerator shown in Figure 4(a), and the cooling capacity at 4.2K in the case of Figure 4(b) also improved to 1.3 times greater than that in the case of Figure 4(a).
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Figure 8 (a) Cooling capacity of 4K GM cryocooler using only Er3Ni and (b) cooling capacity of 4K GM cryocooler using combination of Er3Ni and ErNio.9Co0.a Effect of Regenerator Material on Cooling Capacity The effect of specific heat of regenerator material on cooling capacity was investigated, using Er3Ni and ErNiogCoo i as regenerator materials. Figure 7 shows specific heat of the regenerator materials. Two cases were examined. In case one, experiments were carried out using a regenerator packed only Er3Ni. The other case used a regenerator packed Er3Ni in two thirds of the high temperature side of the regenerator and ErNiogCoo 1 in one third of the low temperature side of the regenerator. In both cases rectifiable meshes were packed in the regenerator as shown in Figure 4(b). The experimental machine employed was the 4K GM cryocooler shown in #5 of Table 1. Figure 8 shows the cooling capacity of the 4K GM cryocooler using only Er3Ni and the cooling capacity of the 4K GM cryocooler using combination of Er3Ni and ErNiogCooa. The cycle frequency was fixed at 42rpm. When only Er3Ni was employed, the cooling capacity was 1.8W, and when the combination of Er3Ni and ErNiogCoo i was employed, the cooling capacity was 2.2W. The cooling capacity of the 4K GM cryocooler using the combination of Er3Ni and ErNiogCoo 1was 1.2 times greater than that of the GM cryocooler using only Er3Ni. CONCLUSIONS ,.,.,..W (1) The cooling capacity of "~ '~ at 4.2K was obtained by the 4K GM cryocoolcr. The main features of this machine are its large size expansion space, the use of the regenerator packed rectifiable meshes, and the combinational use of Er3Ni and ErNio.gCoo.1 as regenerator materials. (2) A regenerator packed rectifiable meshes in two additional places of the regenerator at equal intervals can improve cooling capacity 1.3 times greater than a regenerator packed rectifiable meshes only at the ends of the regenerator. (3) The cooling capacity of the 4K GM cryocooler using the combination of Er3Ni and ErNio.gCoo.1 can be improved 1.2 times greater than that of the 4K GM cryocoolcr using only Er3Ni. REFERENCES 1. 2. 3. 4. 5. 6. 7. 8.
Yoshimura, H.,et al.,Rev.Sci.Instrum. 60(1989) 3533-3536 Inaguchi,T.,et al.,Cryocooler 6 (1990)25-36 Watanabe,K.,et al.,Cryogenics 34,ICEC Suppl. (1994) 639-642 Kuriyama,T.,et al,Cryogenics 34,ICEC Suppl. (1994) 643-646 Yokoyama,S.,et al, MT-14(1996)in press Nagao,M.,et al.,Adv.Cryog.Eng. 39(1994) 1327-1334 Tokai.,Y.,et al.,Jpn.J.Appl.Phys.Partl_ 31 (1992) 3332-3335 Onishi,A.,et al.,Cryogenic Engineering 31 (1996) 162-167
Optimization of Intake and Exhaust Valves for 4 K Gifford-McMahon Cryocooler
Rui Li, Atsushi Onishi, Toshimi Satoh, Yoshiaki Kanazawa R & D Center, Sumitomo Heavy Industries, Ltd., 63-30, Yuhigaoka, Hiratsuka, Kanagawa, 254, JAPAN
The influence of intake/exhaust valve timing on performance of 4 K GM cryocooler has been investigated at various cooling temperatures in order to optimize the intake and exhaust valves. The 4 K G M cryocooler employed for this study is a standard two-stage type one with a rotary valve for intake and exhaust. The second stage regenerator in the cryocooler has both lead spheres and spherical magnetic regenerator material, ErNi0.gCo0.,, with the latter arranged at the colder end. The cooling capacities of the first stage and the second stage were measured in wide temperature ranges (-< 80 K for the first stage, - < 1 0 K for the second stage). The results indicate that the optimum intake/exhaust valve timing is dependent on the cooling temperature and is possible to be different for each stage.
INTRODUCTION In the last several years, regenerative cycle cryocoolers, such as G ifford-McMahon (GM) cryocooler especially, developed remarkably because of the successful use of magnetic regenerator materials. A substantial body of literature on applying magnetic regenerator materials to the cryocoolers has been published[l], but the papers describing the improvement in the components other than regenerator material are very few[2,3]. As a matter of fact, 4 K GM cryocoolers are developed from commercially available GM cryocoolers, and the optimization of the other components for the cooling at 4 K is as important as the advances in regenerator materials. Intake and exhaust valves are key components of GM cryocooler, and include several important factors, such as valve type, valve structure and valve timing. In the previous researches[4,5], we investigated the influence of valve timing on performance of 4 K GM cryocooler, and reported that a reasonably early open timing of intake/exhaust valve not only brings about a cooling capacity more than 1 W at 4.2 K on the second stage but also produces a much large cooling capacity at the first stage. References of [4] and [5], however, mainly focused on the cooling capacities at 40 K for the first stage and at 4.2 K for the second stage. In order to optimize the intake and exhaust valves at various cooling temperatures for 4 K GM cryocooler, we measured the cooling capacities of both the stages in wide temperature ranges ( - < 80 K for the first stage, =0.8. Certainly, too small NTU (e.g.NTU.. F---
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~9 The design goal (not measured) Refrigerant is helium gas, and the charged pressure was determined 1.3MPa finally as an optimum value for this cooler by experiments. We have chosen the type of the dual opposed pistons so as to eliminate vibrations. The operating frequency was settled at 50Hz. The stiffness of the mechanical spring of the piston has been determined 1500N/m by experiments. We employed dry polytetrafluoroethylene (PTFE) bearings and clearance seal technology for each piston. Moving mass design allows for a very simple, light weight and compact motor which requires no additional mechanisms or guides to support a moving mass. In order to use the narrow gap space efficiently we employed no bobbin coils as the driving coils. The stroke of the individual piston is 5mm at the nominal input power of 8W, and the allowable maximum input power is 15W where the dual opposed pistons don't collide each other.The inner diameter of the connecting tube is 2mm, and are connected to the compressor. We contrived a simple construction at this part and achieved to make a light weight cooler using a copper gasket at this part. We employs 316L Stainless Steal for expander cylinder. Its thickness is 0.1mm, it is mechanically processed to a minimum as possible so as to reduce the heat conduction loss. The cold head which is made of copper, the cylinder and the fin are brazed in vacuum.The stainless steel parts are TIG welded. The displacer is made of epoxy resin, and the regenerator mesh is #180 and its line diameter is 30~m. The number of regenerator mesh is about 1200. We employed dry PTFE bearings for the displacer. The specific frequency in displacer mass-spring system is 1.4times as large as the operating frequency. Compressor Unit Electrical connecting pins ~~
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(Dimensions in mm)
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EXPERIMENTAL PROCEDURE Schematic diagram of measuring system of a cooling capacities is shown in Figure 2. At the cold head, a load heater and a C-C thermocouples were attached to the cold head. The expander was covered by super insulation in order to reduce ~ e radiation heat loss, and was mounted to a dewar. The pressure in the dewar was kept less than lxl0-"Torr by using a turbo molecular pump. All experiments were performed in a room airconditioned at 293K. The measurements of the cooling capacity was made by measuring the input power applied to the loading heater while the temperature of the cold head was kept constant. The measurements of the input power applied to the cooler was made by taking indicated values of the power meter. To avoid the contamination of refrigerant, the parts of the compressor unit and the expander unit were baked at 373K for 72 hours in vacuum. After assembling the cooler, it was baked at 373K for 24 hours in vacuum. After this the cooler was charged by helium gas. Compressor u n i t A.C. Power source I PCR500 / I
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Figure 2. Schematic diagram of the measuring system of cooling capacities
EXPERIMENTAL RESULTS Figure 3 shows the measured cooling capacities of the proto type cooler on three different cold head temperatures at the ambient temperature of 293K. At the cold head temperature of 80K, the proto type cooler had a cooling capacity of 210mW over the design goal when the input power was nominal. And a maximum cooling capacity at cold head temperature of 80K was above 400mW when the input power was maximum. The cool down time from the room temperature of 293K to 80K was about 6 minutes and the ultimate cold head temperature was 51K without loading. The cooling capacities of 10 expanders were measured connecting with a same compressor. Figure 4 shows the results. All measured values had a good reproducibility. Photo 1 shows the proto type micro Stirling cooler. SUMMARY 1. We have developed a micro Stirling cooler, which is 420g in weight and 250cm 3 in volume for an application of a portable infrared camera. 2. We have chosen a split type and a linear drive compressor with dual opposed pistons which we considered to be fittable for long life. 3. The cool down time from 293K to 80K was about 6 minutes, and the ultimate cold head temperature was 51K, and the cooling capacity at nominal input power of 8W was 210mW. 4. We will evaluate the life of the cooler.
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Computer-Aided Design of Split-type Stirling Refrigerator B.J.Huang and H.Y.Chen Department of Mechanical Engineering, National Taiwan University, TAIWAN S.B.Chien and T.F. Shieh Chung-Shen Institute of Science and Technology, TAIWAN The design of split-type Stirling refrigerator mainly relies on engineering experience. In the present study a computer-aided design tool "STC15b" which can be run on PC was developed for the design of split-type flee-displacer Stirling refrigerator. The analytical solution resulted from a linear network model was used in the performance calculation. An empirical relation for the displacer loss coefficient Ca was obtained. The cold-end temperature prediction at various cooling capacity and operating frequency is shown very accurate.
INTRODUCTION
The design of split-type flee-displacer Stifling refrigerator mainly relies on engineering experience. Many researchers attempt to develop a computer-aided design tool using various analysis. The achievement is however quite limited since a complicated numerical method is always involved and the computation requires a super computer for solving the transient governing equations of the various parts of the Stirling refrigerator. In the present study we develop a computer-aided design tool which can be run on PC. The analytical solution resulted from a linear network model was used in the performance calculation. The computation is thus fast and can be done on a PC. To improve the accuracy, the analytical errors due to the unmodeled factors are corrected by an empirical relation collected from field tests. LINEAR N E T W O R K MODEL OF SPLIT-TYPE STIRLING REFRIGERATOR
Stirling refrigerator operates at a cyclic state and the physical process is thus transient. A linear network model is thus developed in the previous study [1] for the system analysis of split-type Stirling refrigerator (Figure 1). The linearly-perturbed models of the components are derived from the governing equations through linearization and approximations. Connecting the equivalent circuits of the components according to the process of Stirling refrigerator, we obtain an equivalent network of split-type Stifling refrigerator as shown in Figure 2. Block diagrams are drawn for the connecting tube and the regenerator in Figure 2 since they belong to the distributed systems. Solving the linear dynamics equations, we obtain two transfer functions of the Stirling refrigerator: Xd (s) [ G3(s)G8(s)/G7(s)+G2(s)][Imp(s)Zti+Imm(S) l+ZtJZ~ ][ sp~R 7~ Ca (s) - (1) Xp (~) Pe (s) a2(s)a6(s']} G ep(S) = --k/G4(S)+ 1-G5(s) G7(s) + 1-Gs(s) • Xp (~) (2) l +Zti/Zc X R Tc G8(s)+G7(s)Zcttanh[F tLt] . where Zti = G7(s)+[G8(s)/Zct]tanh[FtLt]' G1 (s) = Dew(S)+Dde(S)Rpp(S); G2(s) = Dde(S)Rpm(S); G3(s) = G1 (s) + G2(s)Wmp(S); G4(s) - Rpp(S) + R(s)mmp(S); G5(s) - G2(s)mmd(S); G6(s) - Rpm(S)mmd(S); G7(s)ptoS - G8(s) mto (s); Wmp--sVwo/(R ~;'w); Wmd--spoAaw/(R [I'w); Z~ - R T~ /(sV~o); 385
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[mmd(S)G2(S)]- Eme(S) [ Rpm(S) + G2(s)G6(s)]
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It is found that the piston displacement Xp is the system input of a split-type Stirling refrigerator. The system outputs are the displacer displacement Xd and the expansion space pressure Pe. The Stirling refrigerator thus belongs to a single-input-multiple-output (SIMO) dynamical system. The maximum available cooling capacity can be evaluated by integrating the pressure peand volume Ve of the expansion space, where V~ is related to Xd. Since the piston motion as well as the associated pressure waves approaches sinusoidal, Xa and pe can be computed simply from the gain and phase of the frequency response functions, Gdp(jO~) and G~p(jO~). Assuming that adp is the gain of Gdp(jO~) with phase ~d leading the piston; aep is the gain of G~p(jO~)with phase ~e leading the piston. Then, the maximum available cooling capacity is
Qmax- f ~2'~pedV~ =fxa,paapXRpoA aeSin(d~d- dpe)
(3)
where Xpo is the amplitude of the piston; f is the operating frequency; ~ is the piston angle; P,o and Xao are the amplitudes of pe(t) and Xa(t), respectively. The net cooling capacity Qnet c a n be evaluated by subtracting the heat losses from the maximum available cooling capacity Qmax. There are four types of heat losses: namely, heat conduction loss of regenerator Q~o,,a , enthalpy flow loss of regenerator Oenth , shuttle heat loss of displacer Q,h,ttz~ , and hysteresis loss of spring Wir [2].The heat loss due to the gas leakage through the clearance between the displacer and the cylinder wall is related to the seal design, manufacturing process, and material used. In the present analysis, a loss coefficient Cd is included. Cd is the coefficient accounting for the friction and gas leakage of the displacer. Since the theoretical prediction of Cd is very difficult, Cd is analyzed from test results. COMPUTER-AIDED DESIGN TOOL - - T H E PACKAGE "STCI5b"
A computer package "STC15b" was developed in the present study. The package whose configuration is shown in Figure 3 includes three parts: subroutine STC15A, STCCd and STC15s. STC15A is used to evaluate Cdfrom the test results; STCCd is used to derive an empirical correlation for Cd, i.e. Cd(f, TL); STCs is used for the practical design using the empirical relation Cd(f, TL). For design purpose, the subroutine STCs is run to calculate the system performance. After the hardware tests, the subroutines STC15A and STCCd will be run to update the empirical relation Cd(f, TL) in order to obtain a better result than before. The package "STC15b" is run on PC. The computation time takes a few seconds for each design. RESULTS
Some split-type Stirling refrigerators were designed and built in the laboratory for experiments. An empirical relation was obtained: Cd(f, TL): F(TL) • G(/) (4) where F(TL) - 430.66 + 1.0851 T,~ + 77346•176 T~ ' 9G(]) - 0.6653 +0.2604f- 5.212 • 10-4ja The design of a Stifling refrigerator using the package "STC15b" can then be carried out. Shown in Figure 4 is the PV diagram of the expansion space computed from the package.; Figure 5 is the cooling capacity versus frequency; Figure 6 is the cooling capacity variation with the cold-end temperature. The prediction of cooling capacity using "STC15b" is satisfactory as shown in Figure 6. It can be seen from Table 1 that the errors of the cold-end temperature prediction at various cooling load are mostly within_+_+3K.
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CONCLUSION It is shown in the present study that the design process of a split-type Stirling refrigeratoris can be greatly simplified by using the software package "STC15b" The package is also designed to have a learning capability in order to accumulate the field experience and update the program automatically.
Acknowledgment
The present study was supported by the National Science Council, Taiwan, through Grant No. CS84-0210-D-002-027.
REFERENCES 1.Huang B.J. and Lu,C.W. Linear network analysis of split-type Stirling refrigerator". Cryogenics (1994)34 ICEC Supplement 207-210 2.Urieli, I. and Berchowitz, D . M . "Stirling Cycle Engine Analysis". Adam Hilber Ltd., Bristol, UK. 1984 T a b l e 1 C o m p a r i s o n b e t w e e n test results and the design calculation using " S T C 1 5 b " .
frequency, Hz QL= 0W 25 30 35 40 45 QL= 0.2W 25 30 35 40 45 QL-~ 0 . 4 W . 25 30 35 40 45 . . . .
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Cryocoolers
Other coolers
This Page Intentionally Left Blank
Microcooling of Low Temperature Electronics
H.J.M. ter Brake, J.F. Burger and H. Rogalla University of Twente, Faculty of Applied Physics, PO Box 217, 7500 AE Enschede, The Netherlands
Small scale low temperature electronic applications would largely benefit from the development of a closed-cycle microcooler, with a cooling power in the milliwatt range. We started a microcooling project, and the first cooler configuration under study is the Joule Thomson cycle in combination with sorption compressors. A sorption compressor set-up was built, and a thermodynamic model was developed. Model and experimental results were in good agreement and can be used for future research.
INTRODUCTION Low temperatures provide an excellent operating environment for conventional and superconducting electronics. For conventional electronics, it increases the speed of digital systems, it improves the signal to noise ratio and the bandwidth of analog devices and sensors, and it reduces aging. For superconducting electronics, it is an essential operating condition. There is a broad range of applications making use of cold electronics, such as computers, amplifiers, mixers, fast AD and DA converters, IR detectors and SQUID magnetometers [1]. Many of these systems need only a modest cooling power, as low as 10 mW or even lower. Standard cryocoolers are largely oversized for such devices, and the development of a microcooler would be of great use for low temperature electronics. Walker [ 1] has given an overview of small scale cryocoolers. Little [2] scaled down the conventional Joule Thomson cold stage using photolithographic techniques for patterning gas channels in glass layers. This cooler, however, has the disadvantage that high pressure gas is needed from a relatively large storage bottle. To our knowledge, the smallest closed cycle cooler realised so far is manufactured by Inframetics [3]. They make integral Stirling coolers of 0.3 kg weight, 9 cm maximum size, and 0.15 W cooling power at 80 K with an input of 3 W. A further reduction in size will be limited by the manufacturing techniques and in this respect micromechanical techniques have been suggested as an alternative [4]. We recently started a project with the aim to realise a closed-cycle microcooler. The project is carried out in cooperation with the MESA Research Institute because of their experience in micromechanical technologies. These technologies use special deposition, etching and bonding techniques - originating from the IC,technologywith which very small mechanical structures can be made, in or on planar substrates (mainly silicon and glass). A typical example of a microchannel structure is given in figure 1 [5]. This paper describes the microcooling project. The main requirements and some design considerations for microcooling are given: After that, a sorption cooler is described, and modelling of a sorption compressor is compared with test experiments. Figure 1 Deep anisotropic etching in silicon [5]. This research is supported by the Dutch Technology Foundation (STW). 391
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REQUIREMENTS FOR MICROCOOLING Requirements for microcooling of low temperature electronics can be summarised as follows: 1. Working temperature: 60- 80 K. Many HTc superconducting devices function at LN2 temperature. The temperature of conventional electronics is not very critical. In general, it should not be lower than 40 K due to carrier freeze-out below that temperature. Therefore, LN2 temperature is very suitable for many applications, also for hybrid electronics. 2. Cooling power (net): 5 - 50 mW. The required cooling power is determined by three factors: the power dissipated in the cold device, the heat leakage from the environment and the desired cool-down time. The power dissipated by superconducting electronics is very small, even a microprocessor with thousands of junctions on a substrate of 5 mm x 5 mm may dissipate less than a few milliwatts [6]. Dissipation of conventional electronics is more severe, therefore only small circuitry can be cooled. The heat leakage is determined by conduction through the supporting structure and the wires, and by radiation. Conduction through the wires does not play a major role. Radiation at 60 - 80 K can be reduced to roughly 1 mW via appropriate shielding. The cool down time is determined by the combination heat capacity - cooling power. Furthermore, the required gross cooling power will be much higher in order to deal with losses by conduction through the cooler itself. 3. Closed cycle with long life-time. We aim to realise a life-time of several years. This puts restrictions on the use of moving components, especially if friction is present. 4. Low electromagnetic and mechanical interference. 5. Other issues: fast cool down time, low heat rejection (i.e. high efficiency), small size, low price.
DESIGN CONSIDERATIONS Commonly used cryocoolers can be divided in recuperative (JT, Brayton) and regenerative (Stirling, GM, pulse tube) cycles. Most of these cycles use the reversible expansion of an ideal gas, performing work on the environment and taking up heat at low temperature. Only Joule Thomson expansion is an irreversible process, performing internal work on a non-ideal gas (no moving parts at the cold side). Interesting alternative cycles worth considering are: thermo-electric cooling and the desorption of a gas from a solution. In the field of micromechanics much experience has been obtained in the design and realisation of fluid and gas handling systems. Also, all kinds of elastic membranes and actuation principles have been realised. An important design constraint is that rubbing surfaces, occurring in nonflexible constructions such as pistons and hinge points, are highly undesirable in microtechnology. Some considerations with respect to the miniaturisation of cooling cycles" 1. Stirling. Miniaturised membranes may be an attractive alternative for the use of pistons, but the relatively low pressure differences require highly efficient regenerators. Interferences are likely to occur at the cold side of a Stirling cycle. Bowman et al. [7] recently patented an idea for a microminiature Stirling cooler with membranes running at a high frequency (500 Hz or more). 2. Pulse tube. This cycle is very interesting because no moving parts are used at the cold side. Membrane actuation may be possible. A patent of Cabanel et al. exists on a micro pulse tube [8]. 3. Brayton. A very small turbo expander requires a very high rotation speed, and seems not feasible. 4. Joule Thomson. Little [3] has proved that miniaturisation is possible. However, a small reliable compressor being able to generate high pressure differences is a major requirement for closed cycle operation. A possible candidate for this is a sorption compressor, and the microcooling project therefore started with research in this field.
SORPTION COOLER Basic operation Figure 2 gives the set-up of a sorption cooler [9]. It consists of a compressor unit, a counterflow heat exchanger, and a Joule Thomson expansion valve. The compressor unit contains four sorption compressors and several check valves to control the gas flows. Figure 3 gives a sketch of a sorption
ICEC16/ICMC Proceedings ]1
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compressor. Low and high pressures are generated by cyclic ad- and desorption of a working gas on a sorption material, which is accomplished by cooling and heating of the sorption material. Cooling is done with a gas gap heat-switch between the sorption material and a cold source on the outside. A compressor cycle is schematically drawn in figure 4. The compressor is heated during sections A and B, and cooled during C and D. During sections A and C both valves of the compressor are closed, and the compressor is in a regenerating phase. During sections B and D one of the valves is opened; the compressor generates a high pressure gas flow during B, and a low pressure gas flow into the compressor during D. Except for some check valves, the cooler has no moving parts in it. This minimises electromagnetic and mechanical interference, and it contributes to a long life time. Moreover, distributed spot cooling can be realised very well with several small micromachined JT coolers. Important aspects in the development of a very small sorption cooler are: the selection of gas/sorber combinations that are efficient around ambient temperature, the choice of a gas or a gas mixture with a high JT efficiency [ 10], the development of small compressors with an efficient heat management, and the development of small micromachined high pressure check valves. Compressor model and test experiments For predicting the behaviour of a miniaturised sorption compressor, a thermodynamic model for a cylindrical compressor was developed, as well as a test set-up for model validation and gas/sorber research. The compressor consists of five elements: heater, sorption material, inner container, gas gap and outer container - see figure 3. In the model the heat flows between these elements are calculated and used to calculate the temperatures, as a function of time. Also a radial temperature profile in the sorbent is calculated. At each time step the amount of gas adsorbed in relation to that temperature is calculated (using Polanyi's theory [ 11 ]), and by that the pressure in the compressor can be predicted. With the test set-up the following parameters can be determined: the temperatures at four positions in the compressor (see figure 3), the amount of gas that flows into the compressor (and thus the total amount of gas in the compressor), and the pressure in the compressor. The heating rate and the temperature of the heater can also be controlled. Firstly, the test set-up was used to characterise the adsorption of nitrogen on activated charcoal. The theory of Polanyi was confirmed, only at low pressures in combination with low temperatures theory and experiments did not correspond. Secondly, the time dependence of the temperatures and the pressure in the compressor were measured. In figures 5a and b typical results are compared with the model. In this experiment the compressor was filled with a measured amount of gas and cycled from room temperature to a temperature set point of the heater, and back again to room temperature. A good correspondence between the model and the experiments was found. Model and experiments are discussed in more detail elsewhere [12].
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Sorption cooler feasibility Most important for the feasibility of a sorption cooler is an efficient and suitable gas/sorber combination. When such a combination is found, scaling down of the compressor seems feasible. The heat stored in the compressor should be as small as possible, since it is parasitic heat. The heat capacity is determined by that of the sorber and that of the container material, and it can be shown that the ratio of those two is independent of the compressor size. Therefore, the heat required for warming up the compressor scales linearly with its volume. In first order, when the cycle frequency of the compressor is kept constant, the input power and the resulting mass flow (with which the cooling power is proportional) scale linearly with the compressor volume (implying a constant efficiency). Also, essential aspects for the miniaturisation of a sorption compressor are: good thermal conduction within the compressor (from heater to sorption material and from sorption material to heat-switch), and miniaturisation o f the heat switch.
CONCLUSIONS Research was started towards miniaturisation of closed cycle cooling methods. Application of micromechanical techniques for components of a microcooler is also investigated. A miniaturised sorption cooler with JT expansion is a promising candidate for small scale cooling. The developed model of a sorption compressor was in good agreement with measurements. Miniaturisation is feasible, but the availability of a gas/sorber combination that is efficient at room temperature is a crucial requirement. In the future, attention will also be paid to miniaturised check valves, a miniaturised heat switch, and the possibility of miniaturising regenerative cooling cycles.
REFERENCES 1 2 3 4 5
6 7 8 9 10 11 12
Walker, G., Low-capacity cryogenic refrigeration, Oxford University Press (1994) Little, W.A., Microminiature Refrigeration, Rev. Sci. Instrum. 55 (5) (1984) 661-680 Inframetics, Inc., 16 Esquire Road, North Billerica, MA 01862-2598, U.S. Walker, G. and Bingham, R., Micro and nanno cryocoolers: speculation on future development, 6 th Int. Cryocooler Conf. (1990) 363-375 Jansen, H.V., de Boer, M., Burger, J.F., Legtenberg, R., and Elwenspoek, M., The black silicon method II: the effect of mask material and loading on the reactive ion etching of deep silicon trenches, Proc. Workshop on Micro and Nano Eng., Davos, Switzerland, 1994. Kirschman, R.K., Low-Temperature electronics, IEEE Circuits and Devices, 6 (2) (1990) 12-14 Bowman, L., Berchowitz, D.M., Urieli, I., Microminiature Stirling cycle cryocoolers and engines, U.S. Patent 5 457 956 (1995) Cabanel, R., Friederich, A., Refroidiseur ~agaz puls6, demande de brevet european No. 0672 873 A 1 (1995) Wade, L.A., An overview of the development of sorption refrigeration, Adv. in Cryogenic Eng. 37 (1992) 1095-1106 Alfeev, V.N., Brodyansky, V.M., Yagodin, V.M., Nikolski, V.A., Ivatsov, A.V., Refrigerant for a cryogenic throttling unit, UK Patent 1 336 892 (1973) Polanyi, M., Verhandl. Deutsch. Physik., Vol. 16 (1912) Huinink, S.A.J., Burger, J.F., Holland, H.J., van der Sar, E.G., Gardeniers, H., ter Brake, H.J.M., Rogalla, H., Experiments on a charcoal/nitrogen sorption compressor and model considerations, 9th Cryocooler Conf. (1996)
Low Cost Mixture Joule Thomson Refrigerator
A.Alexeev, H.Quack, Ch.Haberstroh Technische Universit~it Dresden, Lehrstuhl ftir K~ilte- und Kryotechnik, D-01062 Dresden, Germany
A closed cycle Joule Thomson Refrigerator for cooling electronic devices, which operates in the temperature range from 70 K to 95 K, has been developed which uses refrigerant mixtures and a single stage oil lubricated compressor. It provides high reliability and no maintenance. This system is compact, has a good thermodynamic efficiency and low levels of vibration and noise. The system and its performance characteristics will be described in this paper.
INTRODUCTION Many future applications of HTSC will be in high frequency electronics techniques. Nowadays thin layer SQUIDs made from HTSC are under discussion for a variety of applications. These include in medical research the measurement of biomagnetic fields (human heart signals etc.) as well as in material science the non destructive testing of sensitive constructions or in geological survey the exploration of the earth natural resources. A cooler should not interfere with the measurement taken and it should have a high reliability. Also handiness and mobility are a necessity for out-of-laboratory applications. The state of the art to this date did not provide a sufficiently good solution. Thanks to recent advances in the development of mixture Joule Thomson refrigerators, this deadlock seems now to be overcome. The commercial application of reliable coolers is within reach to costs which compare to those of the overall system. BACKGROUND The use of multicomponent refrigerants in cycles of Joule-Thomson type in the range of temperatures T>100 K has been known for a long time (W.J.Podbelniak, A.P.Klimenko, M.Fuderer and A.Andrija, D.J.Missimer). Using mixtures in systems, working at temperatures T~TII-TI, was observed by the operations, and expressed in Figure 4 with N. Here we find a saturation tendency and
~TN=I.2K at N=6.
In order to analyze this character, the time variation of TI,(N=6) from the level at t=0 was computationally simulated with an ideal condition of no back flow through the thermal diode and no heat income from the atmosphere, and the result is shown in Figure 5.
Substantially different from the
observed T(obs), the computed T(ideal) shows large decreasing without any saturation. However, when the heat back flow is taken into account by a condition of remaining ethanol film of 80 lain thickness on the
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condenser plate surface, the computed T(back tlow) shows large suppression of the lowering. Still existing difference from T(obs) will be attributed to heat leak from the atmosphere in vacuum by radiation and also conduction via support contact. This heat income per each step operation is proportional to the period ~t, and then more frequent operation with shorter period is required for obtaining low temperature and large heat flux pumping, while the pericxt is limited by the thermal relaxation time "~. Accordingly, an improved construction of the (Gd+HP) unit was designed as depicted in Fig.2(b), where the Gd specimen was cut into sliced plates and incorporated in the heat pipe. This structure will be effective to diminish the heat resistance inside and outside of the Gd substance. Three units of this combined (Gd+HP) system are set in a serial connection as shown in Fig.2(b). RESULT The step magnetization pr(xzess was carried on this combined system and the temperature changes from t=0 are presented in Fig.6, where we can find as first remarkable improvement in the heat relaxation time shortening by the stepping period of ~t in sec. The next point to be noted is the different character in time variation of Tt! and TI~. After rapid cooling due to the demagnetization of the adjacent unit (see Fig.2(b)), TIt starts to increase by the forward heat tlow from the central unit magnetization, while TI~ continues to decrease with little counter flow. This indicates a real evidence of the heat rectifying operation by asymmetrical heat pipe structure. The Tt~ and TI. changes were well expressed by exp[-t/~] and the thermal relaxation time -r through the 3 units was determined at T(lbllow)= lmin. with -r(counter)/l:(follow)>_ 6. This shorter ~ value shows remarkable improvement in the heat conductance of about 6 times to the Gd+HP separated system of Figure 2(a) and we can expect significant progress in a new type magnetic refrigerator designed by use of this incorporated unit of (Gd+HP) type. REFERENCES 1 Barclay, J.A., Magnetic refrigeration Adv. Cry. Eng. (1988) 33 719-731 Barclay, J.A, A review of magnetic heat pump technology Proc. 25th Intesoc. Energy. Convers. Eng. Conf. , USA (1990) 7 222-225 2 Nakagome, H. and Tanji, T. and Horigami, H. and Numazawa, T. and Watanabe, Y. and Hashimoto,T., Adv. Cry. Eng. (1984) 29 581 3 Brown, G.V., Magnetic heat pumping near rcx~m temperature J. Appl. Phys. (1976) 47 3673-3680 4 Aoki, R., Magnetic refrigerator operating from room temperature with permanent magnet Cryogenic En~.(JaDan) (1985) 20 294-301 50chi,
T. and Aoki, R., Proc. 10th Intl.Workshop on Rare Earth Magnets and theirAppl. Kyoto, Japan
(1989) 195-201 6 0 c h i , T. and Ogushi,T. and Aoki, R., Proc. Intl. Workshop on Thermal Invest. of ICs and Microstr. Grenoble, France (1995) 150-153 7 0 c h i , T. and Ogushi, T. and Aoki, R., Development of a heat pipe thermal diode and its heat transport performance JSME Intl. J. (1996) 39B No2(in print) 8 0 c h i , T. and Aoki, R., Magnetic c(x~ling and refrigerating effect by rare earth element Rare Earths (Japan) ( 1991) No 19 19-30
402
ICEC16/ICMC Proceedings _T L
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CONCLUSION With an actively cooled Cryostat Vacuum Vessel and equipped with a radiator a hybrid SCHeMe IIcryostat can be realized which provides the necessary quiet environment at < 2 K for the scientific payload of STEP while fulfilling stringent mass and volume requirements and providing a lifetime of over 6 months with 20 % margin.
Design of Thermal Shield for the ITER Cryostat
Kazuya Hamada, Kazuhiko Nishida, Takashi Kato, Tadaaki Honda, Hiroshi Tsuji, Akira Itoh, Masahiro Nakahira, Susumu Shimamoto, Michael E.P Wykes*, Robert Bourque*, Isamu Ohno** and Yasuyuki Miyauchi** Japan Atomic Energy Research Institute, Naka Fusion Research Establishment, 801-1, Muko-yama, Nakamachi, Naka-gun, Ibaraki, 311-01, Japan, *ITER Joint Central Team, Naka Joint Work Site, 801-1, Muko-yama, Naka-machi, Naka-gun, Ibaraki, 311-01, Japan, **Ishikawajima Harima Heavy Industry Co., Ltd., 6-2, Marunouchi, 1-chome, Chiyodaku, Tokyo, 100, Japan
Japan Atomic Energy Research Institute is studying a thermal shield in a cryostat, whose diameter is 36m and height is 34m, for International Thermonuclear Experimental Reactor. The thermal shield is cooled by 80-K helium gas. The designed radiation heat leak from room temperature to 4-K superconducting coil and the structure is less than 0.24 W/m 2 by using a multi-layer metal plate. Total heat load of thermal shield is 147 kW at 80K. The thermal shield is a segment structure in order to allow a maintenance and replacement by remote handling system.
INTRODUCTION Japan Atomic Energy Research Institute (JAERI) participates in design activities and research and development for the International Thermonuclear Experimental Reactor (ITER) in collaboration with European Union, Russian Federation and United States of America[1 ]. In ITER, superconducting coils, such as Toroidal Field (TF) coils and Poloidal field (PF) coils are used and the total weight is around 25,000 tones. The coils are forced flow cooled by using supercritical helium at 4.5K, 0.6MPa and are installed in the cryostat for vacuum insulation. In order to reduce a radiation and conduction heat load from room temperature to 4-K coil, a inner surface of cryostat is covered by thermal shield system cooled at 80 K. The PF coils are operated at pulse current and induce an electro-magnetic force and a Joule heat caused by the induced current on the thermal shield. The thermal shields are designed to safety withstand these loads and to meet the heat load and cooling condition requirements. Additionally, they are designed to be maintained using remote handling techniques. DESIGN CONDITION AND STRUCTURE The cross section of Cryostat is shown in Figure 1. The outer diameter and the height of cryostat is 36 m and 34m, respectively. Thermal shield are installed over the inner surface of cryostat and around of the many access port to the plasma vacuum vessel, and to coils gravity support. Usage of liquid nitrogen for cooling of thermal shield is avoided because of radioactivation. Therefore the thermal shield is cooled by 80-K, 1.6-MPa helium gas. The requirement and the design criteria for thermal shield are as follows; @Thermal shield is cooled by 80-K helium gas and the inlet temperature/pressure is 80K/1.8MPa and outlet temperature is less than 100K. Pressure drop of overall thermal shield is less than 0.1 MPa. @ Heat flux from thermal shield to 4-K objects is less than 0.24 kW/m 2 @ Heat flux from room temperature to thermal shield is less than 9 W / m 2 for cryostat and 3.6 W / m 2 for 427
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ICEC16/ICMC Proceedings
port and gravity support. @to allow a maintenance and replacement work by remote handing techniques @to withstand thermal contraction and electromagnetic force resulting from pulse operation of PF coils @to withstand an accumulative neutron and gamma dose of 1 X 107 Gy The thermal shield is a segmented structure, as shown in Figure 2. The estimated total surface area of thermal shield is 20,230 m 2. The dimension of one segment is determined as 4 m X l m, to be compatible with remote maintenance and replacement. Each segment is connected by a cover plate which bridges of the gap between neighboring segments so as to block the heat leak from room temperature. On the shield panel, cooling pipe are welded. 36m
._nCryostat v]
J
Segment --" Vertical Port ,~rmalShield
/
PF-2 PF-3 /
I
Horizontal Port
03
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PF-4
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i
,
/
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-I---I-- Exhaust Port
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PF-5 PF-7 i m
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Figure 1
Cross section ofITER cryostat
Figure 2
Thermal shield of upper ceiling
THERMAL DESIGN To accomplish the design requirement of heat flux, a conventional multi-layer aluminum deposited polyester (super insulation) and multi-layer metal plates are compared. The characteristics of each insulation method is shown in Table 1. Table 1
Comparison with insulation method
Multi-layer metal plates Fabrication procedure is complicated Quality control is easy Depend on the work procedure Performance good weak for mechanical property Accumulative Neutron and Gamma On account of superior radiation resistance, multi-layer metal plates structure are selected as thermal Fabrication
Aluminum deposited polyester Available
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shielding for the cryostat. In order to accomplish the design condition of heat flux, the needed number of layer and the available emissivity are estimated as 45-layers and 0.3, respectively. The shield material is a stainless steel plate, which the thickness is 0.1 mm, and aluminum is deposited. The schematic diagram of multi-layer metal shield is shown in Figure 3. The thickness of shield is 95.4 mm and the shield is supported by Titanium rod. The static heat load and the helium gas mass flow rate are calculated and listed in Table 2. Total heat load is 147 kW and removed by 1427.2-g/s helium gas flow. A block flow diagram of cooling path is shown in Figure 4. The cooling path is distributed for 3 cooling objects, such as cryostat wall,, port and gravity support. Each cooling path consists of 4 blocks for circumferentially. Such a paralleled configuration results in a low pressure drop for cooling down operation and small temperature difference within thermal shield. The pressure drop is estimated as less than 0.1 MPa at 80K. It is assumed that the friction factor is expressed as Blasius formula. S h i e l d Support
q5 5 m m P o l y m i d e rod
(q5 1 0 m m T i t a n i u m )
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F igure 3 Table 2
Schematic diagram of cross section of thermal shield (from vertical direction) Static heat load condition of thermal shield Part of thermal shields Cryostat Port
Gravity Support of Coil Total
Heat load of thermal shield 41.6 91.1 14.3 147
Mass flow rate 404 g/s 885 g/s 138 g/s 1427 g/s
kW kW kW kW
Cryostat Upper Ceiling Cover h Cryostat Side wall Cryostat Bottom Cover
180-K gas supply system I
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Block diagram of cooling path per one block (45 ~)
~-~
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ICEC16/ICMC Proceedings
The helium gas is supplied by warm helium circulation compressors located at room temperature and the gas is cooled by liquid nitrogen. W E I G H T LOAD CONDITION The weight of one segment of shield panel is estimated as 270 kg and total weight of thermal shield is estimated as 1,370 tons. 0.2 G is assumed as seismic load condition. The PF coil is operated by pulse .current and the most serious condition for the thermal shield is a discharging operation such as a quench. The most adverse electro-magnetic force occurs at the location induced in Figure 1, during discharge following coil quench. At the thermal shield panel, nearly 1 T is applied by PF-4 coil and dumped at 15seconds time constant. The calculation results are summarized in Table 3. Table 3
Load condition per one segment of shield
Self-weight 270kg
Seismic load Vertical direction 954. lkg Horizontal direction: 80.1 kg
Electro-magnetic force Vertical direction" 22kg Horizontal direction: 65.4kg
In case of a segments located on the vertical wall of the cryostat, the self weight and seismic load are supported by two Titanium ( qb 10 mm) rods from cryostat wall. For a segment which is located at top and bottom of cryostat, the segment is designed as self-standing. CONCLUSION The conceptual design of thermal shield for ITER cryostat is carried out. It is estimated that the required shield performance is less than 0.24W/m 2 and the thermal shields is designed by conventional technique. ACKNOWLEDGMENT The authors would like to thank Drs. M. encouragement and support on this work. frame work of the ITER EDA agreement. ITER Director, the parties to the ITER EDA
Ohta, T. Nagashima and S. Matsuda for their continuous This report is an account of work for undertaken within the The view of authors do not necessarily reflect those of the agreement, or the International Atomic Energy Agency.
REFERENCE
1 Thome, R., Design & Development of the ITER Magnet System, Cryogenics (1994) Vol.34 ICEC Supplement 39-45
DESIGN OF SUPERFLUID-COOLED CRYOSTAT FOR 1 GHz NMR SPECTROMETER
Akio Sato*, Tsukasa Kiyoshi*, Hitoshi Wada*, Hiroshi Maeda*, Satoshi Itoh** and Yoshio Kawate** * National Research Institute for Metals, 1-2-1 Sengen, Tsukuba 305, Japan ** Kobe Steel, Ltd., Electronics Research Laboratory, 1-5-5 Takatsukadai, Nishi-ku, Kobe 651-22, Japan
The basic design of one gigahertz NMR spectrometer is being carried forward. A magnetic field of more than 23.5 T for this spectrometer will be achieved by the superfluid cooling technology. This paper will describe the basic design for the superfluid helium cryostat for an outer superconducting magnet with a cold bore of about 150 mm. Some technical points have become clear. Safety of the cryostat involving a magnet with a huge stored energy of 50 MJ has been checked. The amount of cryogen in the magnet vessel should be less than 100 L. The consumption rate of 708 cc/hr has been estimated.
INTRODUCTION Tsukuba Magnet Laboratory (TML) of National Research Institute for Metals has started the second Multicore research project for the development of a one gigahertz NMR spectrometer. A magnetic field of more than 23.5 T for this spectrometer is a challenging target, but will be achieved, using a newly developed oxide superconducting coil in a backup field of 21.1 T. The cryogenic system consists of two sections: one is a 4 K pool boiling part for the oxide superconducting insert coil, and the other is a superfluid helium cryostat for an outer superconducting magnet with a field over 21 T in a cold bore of about 150 mm. Superfluid cooling is one of the key technologies necessary to achieve a high field of more than 21 T in such a large bore, effectively increasing the critical current for superconducting wire. The magnet design is being carried out for several types. The final design will depend on superconducting wire development status in the coming few years. We took the following magnet size for granted in the cryostat design. The magnet size was supposed to Nb3Sn SC joint be 1200 mm in diameter and 1500 mm in height. The total weight of the magnet will be 8 tons. The magnet has a peculiarity in its superconducting joints; inner Nb3Sn coils have superconducting joints 400 mm above E E the upper surface of the coils as shown in Figure 1. 0 0 This paper will describe a basic cryostat design for this magnet, and point out problems to be solved for the superfluid helium cryogenic system for an NMR spectrometer.
1
NbTi SC joint
BASIC DESIGN OF CRYOSTAT Pressurized superfluid helium cooling has been adopted to achieve long term operation and cryogen reduction in a magnet cryostat, resulting in cryostat compactness and safety insurance in case of a magnet quench. A helium I vessel that supplies coolant to the heat exchanger for cooling helium II is located in the doughnut space around Nb3Sn superconducting joints as shown in Figure 2 and 3. The helium I vessel capacity is 650 L. 431
SC shim coil
SC coil
Figure 1 Arrangement of the NMR magnet and superconducting joints
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Power leads will be removed after energizing the magnet and current transfer to the Persistent Current Switch (PCS). The PCS will be situated in the helium I vessel, preventing heat Mainterminal loss during the energizing process, whereas protection resistance and diode will be located MainPCS in the helium II vessel. Maindumpresister ~ ~ Shimcoildumpresister The superfluid helium acts as a superconductivity stabilizer especially for bare superconducting wire because of its excellent Helium II vessel heat transfer characteristics. In our case, howi:i: ever, all of the coils are impregnated solenoid, Shim coil SC joint so the magnet is almost in an adiabatic condiMainSCjoint (NbTi)----.T tion in the superfluid helium. The superfluid i helium thus acts as only a temperature stabi[ /lI~'~l Shim coil Maincoil l lizer. The amount of helium II coolant is a kind of trade-off between temperature stabilization and cryostat security. In this design, a helium II vessel coolant capacity of less than Figure 2 Schematic view of the cryostat. 100 L has been selected. The heat capacity of The helium I vessel is located in the doughnut 100 L superfluid helium at 1.8 K is 43.4 kJ/ space around Nb3Sn superconducting joint K, so an accidental decrease of cooling power of one Watt would increase the superfluid bath temperature by only 1.38 mK/min. It means that the heat capacity is large enough for temperature stabilization. 100 L of coolant is very small compared with the volume of the magnet. Reduction of the coolant amount in the helium II vessel to less than 100 L will be achieved by filling some compound in the magnet vessel.
MaincoilSCjoint(Nb3Sn)
HeliumI vessel
ii
termiconashi POSm
i.
_ . .
-~bT54
SAFETY ANALYSIS IN CASE OF QUENCH In the case of a quench the magnet will release its storage energy of about 50 MJ. The storage energy will be consumed as Joule heat loss in the magnet, and the magnet temperature will increase up to about 80 K in a few seconds. The superfluid helium around the magnet will begin to increase its temperature and will then evaporate at saturated temperature in the closed helium II vessel. The only exit of evaporated gas is a cold safety valve connecting to the helium I vessel. The flow rate will achieve to a maximum value of 340 L/s in the case where release pressure is 0.14 MPa. Therefore the maximum flow rate through the cold safety valve is estimated to be 4.3 m/s when the valve diameter is 50 mm. All of the liquid coolant inside the magnet vessel will be changed to gas in 1.6 seconds if helium II coolant is 80 L under film boiling assumption, and the release flow rate will decrease to 267 L/s. In this estimate, a cold safety valve of diameter of 50 mm has sufficient capacity for the relief valve. For redundancy, in case of an accident with the cold safety valve, a rupture disk that will release high pressure gas to adiabatic vacuum space
l_
cb 1850
Figure 3 Overview of the designed cryostat
ICEC16/ICMC Proceedings and a drop-off valve on the outer shell of the cryostat will be fitted.
433
/
Suppoting rod
THERMAL DESIGN
-13I . F,F'b I The superfluid heat exchanger in the magnet vessel cools down the helium under conditions of at4 K shield I ... ..... .~!i mospheric pressure. The superfluid heat exchanger is exhausted by a 200 L/min vacuum pump in a steady state at 1.8 K. An alternative vacuum pump is used when maintenance is required. An auxiliary pump with an exhausting rate / of 6000 L/min will be used in the precooling pro/i cess from 4 K to 1.8 K, resulting in a reduction of 1 | J ] [did d~ "] .... u cooling time to half a day. Coolant supplied from the helium I vessel 80K shield Gas Coold Shield (GCS) through the Joule-Tomson (JT) valve flows in a mist state through the heat exchanger, and evapoFigure 4. Arrangement of the magnet vessel and rates on the inner surface of the heat exchanger to the thermal radiation shield. cool down the outside helium in the magnet vessel. Temperature differences between pressurized helium and saturated helium in the heat exchanger are controlled by the JT valve in a certain value to keep the magnet temperature steady. This temperature difference control method was adopted in the Tohoku Univ. magnets and the stable controllability has been confirmed [ 1,2]. The supporting rods and thermal radiation shields location is shown in Figure 4. There are three types Pumpout of thermal radiation shield including Gas Cooled Shield (GCS). The 80 K- shield is cooled by liquid nitrogen, and the 4K-shield for the 1.8 K magnet 'Q=" vessel is cooled by helium I. The magnet vessel is I t Qc.'2 supported directly from the outer vessel by an FRP supporting rods thermally-anchored at 4 K and 80 K and the intermediate GCS temperature, respectively, considering mechanical stiffness. Thermal flows are shown for this case in Fig~ Q r Qr ure 5. An example of the thermal balance calculation result is summarized in Table 1. Total consumption rate of liquid helium is 708 cc/hr. Liquid ".t.Qo.i helium is mainly consumed in the insert Dewar 9~,2 ii ~ that is designed so as to exchange an oxide superOuter v a c c u m e shell conducting magnet, that would be under developQch Conduction through neck tube (helium) ment even in near future. Qcn Conduction through neck tube (nitrogen) We plan to supply liquid helium at the rate Qcs Conduction throuh supporting rod Qcl Conduction in liquid helium of one 500 L Dewar every month. The consumpQr Radiation Qsv Super-leak of liquid helium tion rate almost satisfies this specification. ::
lIQnI ] H
~J
,9 ~
s 3 '1
.,~o
cs 1
Figure 5 Thermal flow model for calculation. PROBLEMS IN LIQUID HELIUM SUPPLY SYSTEM The liquid helium supply will seriously affect the temperature distribution in the helium I vessel, especially in the bottom part of the vessel where lambda point superfluid helium layer is formed in a steady state. Any temperature increase around the communication channel such as the safety valve will induce an effective cooling power decrease. Any temperature change would affect the persistent current of the superconducting magnet with many superconducting joints and induce a magnet quench in the worst case. Therefore one of the key technologies necessary for a stable long term operation of an NMR spectrometer is a reliable liquid helium supply system. A detailed design of the liquid helium supply line exit and the helium I vessel struc-
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Table 1 An example of the thermal balance calculation result, where OVC and GCS are the Outer Vacuum Can and the Gas Cooled Shield, respectively; LN 2 means thermal shield of liquid nitrogen or the stage of 77K; He I and He II also represent its temperature stage ( See Figure 4 and the thermal flow model of Figure 5). The result was calculated for the case of Figure 4.
Heat Flow Path
OVC
~ LN 2
Radiation 50.6 Supporting Rods 2.04 Neck tube (He) 0.519 Neck tube (N2) 0.229 Insert Dewar 8.20 Feedtrough Communication Channels Total
61.6
LN 2 ~ GCS 0.997 0.162 0.045
LN 2 ~
He I
8.45 xl0 -3
0.0190 0.133 0.0164
He I ~ He II 7 . 5 3 x l 0 -6
2.99 xl0 -3
0.216
0.593
1.80
GCS ~ He I
8 . 4 5 x l 0 -3
9 . 1 2 x l 0 -3
5.76 xl0 .3 0.112
0.394
0.121 ( in Watt)
Liquid helium supplied to the superfluid heat exchanger 5.75 xl 0 .3 g/s 166 cc/hr @ 4.2 K; 1 atm Exhausting rate from the superfluid heat exchanger 90.8 L/min @ 300 K Boil-off rate of liquid helium ( Helium I vessel ) 184 cc/hr Boil-off rate of liquid helium ( Insert Dewar ) 359 cc/hr Total consumption rate of liquid helium 708 cc/hr @ 4.2 K; 1 atm Total consumption ra~e of liquid nitrogen 1450 cc/hr @ 77 K; 1 atm
ture will be carded out in the next stage, avoiding disturbance around the communication channel.
SUMMARY In this basic design some technical points have become clear. The amount of cryogen in the magnet vessel is designed to be less than 100 L, which will ensure cryostat insurance in the case of a magnet quench. Safety of the cryostat involving a magnet with a huge stored energy of 50 MJ has been checked. Precooling of the magnet in the restricted space is a remaining problem. The estimated consumption rate of 708 cc/hr for liquid helium will allow the monthly supply system by one 500 L helium Dewar There still remain other problems to be solved. Precooling of the eight-ton magnet will take more than three weeks. Magnet charging will take a few days. In all processes, including normal operation mode, the cooling system operate automatically and stably. We will continue the cryostat design and research and development works to obtain design data.
REFERENCES Watanabe, K., Noto, K., Muto, Y., Maeda, H., Sato, A., Suzuki, E., and Uchiyama Y., Research and Development for Pressurized He II Cooled Superconducting Magnets Sci. Rep. RITU (1986)A-33 297- 306 Noto, K., Watanabe, K.,, Muto, Y., Sato, A., Horigami, O., Ogiwara, H., Nakamura, S., Suzuki, E., Proc. ICEC- 10 (1984, Helsinki), 181 - 184
A Variable Temperature Cryostat to Measure
Jnoncu(T) of ITER Strands up to 20 Teslas
Jager B.*, Bocquillon A.*, Chaussonnet P.*, Martinez A.*, Nicollet S.*, Serries J.P*. and Vallier J.C.** *Association Euratom-CEA DRFC, CE Cadarache, F 13108 Saint Paul Lez Durance, France **CNRS - High Field Laboratory, BP 166, F38042 Grenoble, France
The determination of the non-copper critical current density Jno,,cu is crucial for the development of Nb3Sn strands for ITER coils. The temperature dependence of J,,onCu in a magnetic field plays a leading part in determining the operating parameters and stability of the ITER coils. We describe the development of a variable temperature cryostat based on a new concept. This new cryostat is able to measure critical currents up to 600 A in fields up to 20 T in a temperature range of 4.2 K to 20 K.
INTRODUCTION The different conductors of the ITER magnets are made of cables of superconducting strands. The ITER specifications on the critical properties of these strands mainly concern the critical current density to be achieved at 4.2 K and 12 T and the level of hysteretic losses. The model predicting the critical current density of Nb3Sn as a function of the field and the temperature has been historically presented by Summers [1]. This model involves several parameters and coefficients which need to be carefully characterised for the practical industrial wires used for ITER. Indications already exist on the important variations which take place from one strand to another and this is not surprising in view of the great variety of processes used 9internal tin, modified Jelly Roll, bronze route. More precisely the behaviour of the strands as a function of the temperature has to be checked for every strand. In fact the operating temperature taking into account the design of the cables is never 4.2 K, the usual test temperature for strand benchmarks but generally temperatures greater than 5 K. The new cryostat presented here must help in obtaining information in this direction. It will be able to measure critical currents up to 600 A in the 4.2 K to 20 K range and in magnetic fields up to 20 T. The design allows to insert the cryostat inside the bore of the hybrid magnet of the CNRS High field laboratory.
PRINCIPLE OF THE CRYOSTAT The difficulty in measuring the critical current of ITER type superconducting strands, in magnetic fields which can reach 20 teslas, is to supply a current of about 600 A to the strand to be tested without temperature modification. To solve this problem, we used a new principle (cf. Figure 1). Two current leads are cooled by helium circulation, their two cold ends are connected by a sample support making up a thermal shunt. The superconducting strand to be tested is wound around this support and thermalized by helium which is at the same temperature. The electrical resistance of the support must be high enough for the current going through it to be insignificant, when the strand to be tested is in superconducting state.
DESCRIPTION Test cryostat This is an insert-cryostat which is placed in an existing 80 K cryostat. The general diagram is shown in Figure 1. It essentially includes a helium tank of about 18 liters (height = 1.08 m, external diameter = 0,1643 m, internal diameter = 0.0761 m). This annular tank is pressurised up to a maximum pressure of 435
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CURRENT SUPPLY
PRESSURE REGULATING SYSTEM (P < 0.29 MPa)
>G
L) [.-.
He 2.5 MPa
PCV2 INSULATING TUBE
HELIUM TANK i
I I
9
I
9 80K SHIELD
CURRENT LEADS
~ - - -
i
CRYOPUMP
EVAPORATOR ................................ HEATER
NDOM SEAL
.... - VACUUM
IIII%11 III
HYBRID MAGNET
Figure 1 General diagram of the new cryostat
0.29 MPa. It therefore allows to perform measurements in supercritical helium. A pressure/backpressure regulator system regulates the pressure to the desired value. The helium extracted through the lower part supplies the two current leads. If liquid helium is used, an evaporator/heater helps to supply the current leads in helium gas at an adjustable temperature. Two control valves, placed at the outlet of the current leads, regulate the temperatures of the two cold ends at the same set point (5 K < set point < 20 K). The stability of the temperature is then better than 0.05 K. The sample holder is connected to the cold ends of the current leads. All of the test cryostat is maintained under vacuum by a cryopump. This pump is made by sticking coconut charcoal grains to the bottom of the helium tank. The lower part of the cryostat is removable so as to allow the sample to be introduced. An indium seal ensures tightness versus vacuum. Current leads We used the same technique as that described before [2]. The main characteristics of these current leads are described in Table 1. They consist of a copper braid cooled by helium. This braid is introduced into an insulating tube. This tube is then introduced into a stainless steel tube ensuring tightness versus the external vacuum. This technique helps to avoid cold insulating electrical breaks on the helium supply tubes. The ends of the braids are soldered in OFHC copper end pieces. The temperature of the cold end of these current leads is regulated by the TCV valves (of. Figure 1), which act on the helium flow rate. The electrical and thermal connection with the sample is ensured by a pinched connection on the cylindrical ends of the current leads. The surfaces in contact are machined with care and gilded. Magnetic field temperature measurements The sample temperature must be precisely known whatever the magnetic field. We used new Cernox sensors (type CX 1030) specially developed by Lakeshore for magnetic field measurements. Three sensors are mounted on the sample holder to verify the possible temperature gradient in the presence of current. Two other sensors are mounted on the cold ends of the current leads and help to ensure their
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Table l Main characteristics of current leads copper wire diameter number of wires A, copper cross section L, length of the cooled braid L I/A at I = 600 A
Electrolytic Tough Pitch - 85 < R R R < 100 0.3 m m 624 44.11 10 -6 m 2 1.925 m 2.62 107 A/m
temperature regulation. We measured the effects of a magnetic field up to 20 teslas on these sensors. These measurements were made at constant temperature in a boiling helium bath at 4.21 K. Figure 2 shows, depending on the magnetic field, the errors in terms of temperature for the five sensors. The m a x i m u m error made at 4.21 K remains inferior to + 0.02 K for a field of about 15 teslas. For a AllenBradley carbon resistor, this error is o f - 0.14 K at 20 teslas and continues to increase. Each sensor is mounted in a calibrated hole with grease. Sample The sample (Figure 3) is connected to the current lead ends by two high-purity copper bus bars (RRR 400) which are as symmetrical as possible to minimise the temperature gradient. The reaction mandrel is a V A M A S titanium mandrel which is also kept for the test without any transfer and possible associated degradation. Two copper ends are soldered on the titanium mandrel for the sample terminations. A thin copper cylinder is added after reaction so as to ensure a very uniform temperature of the mandrel. This piece is not in electrical contact with the strand so as to avoid any low resistance short circuit. Two Cernox sensors are installed at the top and the bottom of the sample. The third one is installed in the very middle of the sample. The insertion of the sensor in front of the central turn of the sample is possible by drilling a transversal channel in the mandrel. Two voltage couples, one across the central turn and the other across the 7 turns monitor the resistive transition of the sample.
0.05
.................... i I
..........
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.......
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high purity CODDerbus bars
i5 Cernox sensors CX 1030 7
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titanium .~~mandrel Cernox sensors
Measurements in heliu at T - Cte = 4.21K
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.
.
.
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Figure 2 Compared effects of the magnetoresistance on the temperature measurement at T = Cte = 4.21 K for Cernox sensors and a carbon resistor.
copper 0I I
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Figure 3 Simplified drawing of sample
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ICEC16/ICMC Proceedings 700
600
.
~ 30
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50
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60 0
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ro
3
4
5
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Distance / m m
Fig. 3 Time dependence of the penetration depth of the freezing front, the -20~ isotherm, the-40~ isotherm and corresponding freezing rate in dependence on tissue depth CONCLUSION A cryoprobe with high cooling capacity using liquid nitrogen has been constructed which can be used for endoscopic, e. g. gynecology applications. REFERENCES Herzog, R., Kryotherapieger~it zur medizinischen Behandlung. Luff- und K~iltetechnik 29(1993)33 H~insgen, H.; Binneberg, A.; Herzog, R.; Schumann, B., Chancen der Kryomedizin in der Minimal Invasiven Therapie. Luff- und K~iltetechnik 31 (1995)26
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Thermosiphon Cooler: A Low Microphonic Cooling System for HTC-Devices; Especially for SQUIDs Armin Binneberg, Hanno Buschmann, Ralf Herzog, Johannes Neubert, Gabriele SpOrl Institut ~ r Lufi- und K~iltetechnik Dresden, gGmbH, FB Kryotechnik, Bertolt-Brecht-Allee 20, D-01309 Dresden, Germany Applications of HTC-devices in high-tech areas require a potentially low microphonic (or free) cooling system. Thermal fluctuations, mechanical vibrations and electromagnetic influences are to be minimized (or equalized). A hybrid cooling system was developed consisting of a small split-Stifling cryocooler as refrigeration source and a thermosiphon. The thermosiphon (condenser and evaporator, separated by capillary tubes) is filled with gaseous nitrogen and closed. The condenser is contacted to the cold head of the Stifling machine. The SQUID is mounted on the evaporator and is not affected by disturbances of machine. A cooling power of 0.3 W is reached at 65 K.
INTRODUCTION Use of HTC-devices in high-frequency techniques or other fields of measuring techniques demands special cooling systems adapted to the corresponding measuring tasks. The cooling system has to meet the following requirements: Temperature range for SQUID cooling: < 70 K Nearly no mechanical vibrations Cooling capacity more than 0.1 W
No use of cryogenic liquids Damped electromagnetic noise
All in all the cooling system has to be a compact and portable one. For realization of these points a split Stifling cryocooler from AEG Infrarot-Module GmbH, Germany was selected, having 1 W at 80 K cooling capacity. The thermosiphon was developed for damping and/or separating the influences of mechanical and electromagnetic nature. Furthermore active and passive damping elements were installed. [ 1] EXPERIMENTALS Methods of Compensation of Disturbances Disturbances arising from the mode of operation of the cooler (mechanical motions of parts and eleetromeehanieal signals caused by these motions) must be damped or equalized. So vibrations from the compressor and the cold head (from the split Stirling machine) are diminished by mounting all-metal vibration dampers in the base plate system (see Figure 1). Moreover, the thermosiphon is made as a ,,soft spring", that means the capillary tubes connecting condenser and evaporator are sot~ and formed as an elbow. Bellows are used as the vacuum jacket of the capillary tubes. The evaporator is placed within the sensor dewar by a three point support and the sensor dewar is fastened on the base plate by vibration dampers, additionally. In order to damp electromagnetic noise, sensor place and cold head are separated from each other. The distance between them is up to 300 mm. Condenser and evaporator are at different horizontal levels. Furthermore, electromagnetic shielding materials can be used either as shields or as constructing materials. 509
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COLD HEADOF SPLIT II COLDHEAD
-~L/q_ FAN
AND THERMOMETER THERMOSIPHONASSEMBLY
NITROGEN STORAGE TANK
GAS T R A N S F E R LINE
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Figure 1 Scheme ofthermosiphon cooler Method of Test For testing the thermosiphon cooler temperatures at condenser and evaporator are measured during cooling down. One gets the relation between temperature and cooling time for a set of parameters. The following process variables can be changed as given: Pressure of nitrogen Set point of temperature Volume of the storage tank Diameter of capillary tubes
(0.2... 0.6) MPa. (65,70) K (200, 300, 400) ml (0.8... 1.8) mm.
During the measurements only one parameter was changed whereas the other ones remained still the same. The typical behaviour of the temperatures within the thermosiphon is shown in Figure 2. For valuation of the thermosiphon cooler the amplitude of oscillation was measured at the sensor place from time to time. [2;3] RESULTS AND DISCUSSION The most simple change of parameters was the variation of the volume of the storage tank. But only for a vessel volume of 200 ml the pressure variation doesn't influence the temperature at sensor place. For bigger vessels and pressures of about 0.2 MPa. there is too much nitrogen within the thermosiphon. The circulation of liquid nitrogen is unstable and therefore the final temperature at the sensor place varies with pressure. In Figure 2 the relationship between the cooling down and the diameter of capillary tubes as well as additional masses for simulating a heat load is represented. The diameter of the capillary tubes influences the oscillating properties of the thermosiphon. Enlarging the diameter also means an increase of mass. In that sense the so called ,,soft spring" becomes harder. In dependence of the position of that mass within the thermosiphon the vibrations from the cold head can be damped or increased. Cooling down of the evaporator becomes longer and the final temperature is enlarged. Additional masses simulate thermal heat load at sensor place as well as the cross over behaviour between the variation of diameter. Furthermore, one gets information about the quality of thermal isolation and thermal contacts between all elements of the thermosiphon. This part of investigation shows the complication of laying out such a cooling system for active and passive damping of disturbances from different sources.
ICEC16/ICMC Proceedings
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300 !~
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104 103 10 2
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2
Figure 3" Time variation of VLD at a distance of 5 m r n from the heater. The abscissa indicates time after the onset of heating. Bath temperature T - 1.7 K, peak heat flux q - 20 W / c m 2, heating time tH = 500 #s. Initial VLD L 0 - 102 and 103 crn -2 for the calculation with non-zero source term.
Lo=106 ......... Lo=104 Lo=102
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103 102 101 0.001 0.01
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Time (s) Figure 2" Time decay of VLD from some initial values calculated from Vinen's equation. Bath temperature T = 1.7 K , initial VLD, Lo - 102, 104 and 106 c m -2.
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ICEC16/ICMC Proceedings 1.75
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Figure 5" Experimental data of time variation of temperature profiles at 5 and 30 rnrn from the heater. REFERENCES 1 Murak~mi, M. and Iwashita, K., Numerical computation of a thermal shock wave in He II, Computers & Fluids (1991) 19 443-451. 2 Vinen, W., Mutual friction in a heat current in liquid helium II III. Theory of the mutual friction, Proc. R. Soc. London (1957) A242 493-515. 3
Fletcher, C., Computational Techniques for Fluid Dynamics, Springer-Verlag (1988), 143-170.
4
Book, D., Boris, J. and Hain, K., Flux-Corrected Transport II: Generalizations of the Method, Jour. Comp. Phys. (1975) 18 248-283.
5
Maynard, J., Determination of thermodynamics of He II from sound-velocity data, Phys. Rev. B (1976) 14 3868-3891.
6
Shimazaki, T. and Muraka.mi, M., Measurement of Characteristic Time of Quantized Vortex Development Using a Thermal Shock Wave, to be presented at ICEC 16 (1996).
7
Childers, R. and Tough, J., Critical Velocities as ~ Test of the Vinen Theory, phys. rev. lett. (1973) 31 911-914.
Heat Transfer From Superconductor Wire To Superfluid Helium Yanzhong Li*, Yezheng Wu*, Yuyuan Wu*, Udo Ruppert**, Ingrid Arend**, Klaus Liaders** *College of Energy and Power Engineering, Xi'an Jiaotong University, 710049, P.R. China **Physics Department, Free University Berlin, D-14195 Berlin, F.R. Germany A systematic research on NbTi/Cu wires (single core and multi-filaments) was conducted in HeI, Hells, and HelIp, separately. The results measured in three baths show a large difference in current sharing zone of a same sample, but a similar wire with insulation layer in Hell baths gives the results like those of bare wire in HeI. It means the insulation of superconductor is the main obstacle of heat transfer. In order to, ::5 >~0.2
>~0.2 0.0
l
0.4
B
D
CurrentI, A Figure 4 U-I curve recorded in HeI
A
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Nbri Sl (70~m) at Z06K 94KPa,Hellp
B
. . . . . .C.
.........................,,,,,,,,,,~
D
CurrentI, A Figure 5 U-I curve recorded in Hell
A
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Quench current Ima x and recovery current I r ( from point A and D in figure 4 or 5) as the function of bath temperature Tb are separately plotted in Figure 6. It is seen that the higher current is obtained in better cooling property bath. In HelIp, current increases monotonously along with Tb decreasing. When temperature goes down to 1.9K, the applied current is over the carrying ability of circuit and the sample is burnt out after quench, so the measuring system cannot be used for the measurement at Tb< 1.9K in HelIp. The wire F54-1.35/0.05 was measured in Hells and HelIp. The recorded curves for bare wire are similar to figure 5, but the curves for lacquered wire are similar to figure 4 like bare wire at HeI. The Ima x and I r for F54-1.35/0.05 wire as the functions of T b are plotted in figure 7, in which the results of lacquered wire are also illustrated. It is seen by comparison, the Ima x is reduced by 15% or so in HelIs, and by 30% (Imax) and 60% (Ir) of reduction in HelIp due to the existence of insulation. The Ima x and I r of lacquered wire have no clear difference in Hells and in HelIp, which means the insulation layer breaks down the direct cooling of liquid to solid. In this case, the cryostability of superconductor cannot be enhanced by only improving the liquid cooling property. More consideration needs to be paid to the heat transfer performance of insulation layer. The related research was reported earlier[3]. 35
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Figure 3 Pressure effect on the peak heat flux density of HeII bath at 2.0K when P>P)~(x: experimental results with 511am diameter RhFe wire).
(b) Noisy fihn boiling Contrary to the noiseless fihn boiling state, there exists severe vaporization - condensation phenomena at the boundary, and the sound of boiling can be heard during experiment. A typical example is that for noisy film boiling phenomena, it can be found in heat transfer experiment in HelIs with a hydrostatic head more than 12 cm. The noisy phenomena will possibly disappear when the hydrostatic head goes much higher. For HeIIs at 2.0K, when HeII is slightly subcooled (for example, with a bath pressure of 0.004 Mpa), the noise from film boiling heat transfer measurement will not be heard. This pressure range, though not clearly defined, is just the transition region for heat transfer in HeII, however a quantitative equation of pressure effect on peak heat flux is difficult to show. Subcooled superfluid helium If HelI is slightly subcooled, and when the pressure P is less than or equal to P)., and P is greater than t~.(7}~)(the saturated vapor pressure at T~0, the total subcooled degree is t ' - t ' s + pgh when the wire is submerged in the depth of h in HeII. So the Clausius-Clapeyron equation gives 7' ATi~a~ . ( 1 ' - 1's + p g h ) (9) A,L
If P - l~s,>> pgh, then the peak heat flux density can be written in the equation as following .~ 20(r~,) 7' 1 qp
= . . . . . . r,, p,,L
(lO)
Even if 9gh tends to zero, APA becomes relatively important, the APA item can still be neglected in comparison to ( P - t~,). The applicability of the heat transfer equation in HelI can be tested at small subcooled degree under the condition of t's < P < P;, by eqn.(10), such as the experiments at 2.0K[7]. The experimental result is 15 W / c m 2, and the theoretical one is 15. I W / c m 2 when P is 0.0045MPa. So the relative error is only about 2 % , at this time O(r 0) is 0.029. Furthermore, the calculated peak heat flux q,, is 16.2 W / c m 2 at P=P)=0.005MPa, the experimental one is 16.6W/cm 2 the relative error is only about 2.5% PEAK HEAT FLUX DENSITY qp IN HEll WHEN P > t'~
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It is seen from the phase diagram of helium, that there are GHe, HeI and Hell layers when a wire is heated when the film boiling is initiated at a pressure of P>Px. When heated, the thin film of HeI will be formed around the wire if the heat flux is big enough. In another words, there will appear T=T~ at the radius of r=r0, however HeI will be vaporized in millisecond time because the heat flux is big enough and HeI around wire will change to GHe, GHe will be wrapped by HeI, HeII respectively. The author has fully discussed the calculation of the critical heat flux about the complex phase boundary in reference [ 1]. The fundamental condition needed to initiate film boiling of HeIIp is that there appear HeIIx (T~ = Tt ) at the position of wire/HeII boundary, and a temperature gradient (r = r0, T = T~;r = 0% T - T8 ) exists. When TB deviates from Tx greatly (T8 - T,t > 0.06K) in general, by integrating the Gorter-Mellink equation from TB to T~., we get
, 2~,'o_______~)IT~ dT
qp =
ro
r, f ( T )
(11)
Because the X-line in the phase diagram is a line nearly vertical to the T axis [when P=P~, T~,=2.172K, when P=0.1MPa, T~=2.163K], the effective heat transfer coefficient f(T) ~ (f(T,P) 1 in reality) varies little from the saturated pressure to 0.1MPa. It can be concluded from eqn.(11) that peak heat flux q~ varies little with bath pressure P from 0.005MPa to 0.1MPa. The experimental results of the peak heat flux q~ at 2.0K using RhFe wire with a diameter of 51~.tm when the bath pressure are 0.005MPa, 0.05MPa, 0.1MPa respectively have proved the deduction (shown in Fig.3). The modified factor ~(r 0) is of big difference to the one in eqn.(4) or (10) because of the influences of the boundary phenomena when the film forms and the transient process. From experimental results, we get q~Qb)=0.0709 and the theoretical line is displayed in Fig.3. The symbol 'x' represents experimental results. Thus we can see that the theory demonstrates accurately the relation between P and q~. CONCLUSION The above studies show that Van-der-waals pressure plays an important role in the heat transfer in HeIIs for a small hydrostatic head; Fountain pressure can be a very interesting pressure term for special experiment, which will influence a lot on the critical heat flux in HeII; Sometimes hydrostatic head, Vander-waals pressure and fountain pressure will combine together to influence heat transfer in HeII. In HeII with a small subcooled degree (P>pgh), the influences of Van-der-waals pressure and hydrostatic pressure on peak heat flux of HeII can be neglected, however the influence of bath pressure is clearly shown; When a HeII bath is subcooled to a large degree ( P> Pz ), bath pressure plays little role in peak heat flux. ACKNOWLEDGMENT This work was supported by Chinese National Natural Science Foundation (contract number:59406010). REFERENCES Wang, R., Peak and recovery heat flux densities in bath of subcooled superfluid helium Cryogenics (1994) 34 983-990 Gradt, Th. , Szi~cs, Z. , Denner, H. D. and Klipping, G. Heat transfer from thin wires to superfluid helium under reduced gravity Adv. Cryog. Eng. (1986) 31 499-504 Li Y. Z., The heat transfer properties of superconducting wire and those with porous coatings in superfluid helium PhD Thesis, Xi'an Jiao Tong University, China (May 1995, in Chinese) Li Y. Z., Wu Y.Y., Arend I., L i~ders K., Ruppert U., Influence of porous coatings on heat transfer in superfluid helium Cryogenics (1994) 34 (suppl.) 301-304 Van Sciver, S. W., Helium Cryogenics Plenum Press, New York, USA (1986) Briantsev, K. A. , Sidyganov V. U . , HeII-Vapour interface stability at high heat flux Cryogenics (1992) 32 (suppl.) 253-256. Wang, R., Time dependent heat transfer to subcooled superfluid helium PhD Thesis, Shanghai Jiao Tong University, China (March 1990, in English)
Kapitza Conductance of Niobium for S. R. F. Cavities A. Boucheffa, M. X. Francois and J. Amrit L. I. M. S. I.- C. N. R. S., B.P. 133, Orsay, 91403, France Cedex The Kapitza conductance is measured at the niobium-helium II interface for temperatures ranging from 1.5 K to 2.1 K, using a new experimental method in which the heat flux is directed from the liquid to the solid. These experiments clearly show the effects of chemically polishing rough surfaces and the influence of surface oxides on the heat transfer process. A comparison with existing measurements and theories are also made.
INTRODUCTION The interest in the thermal boundary conductance (Kapitza conductance) between niobium and helium II arises from the fact that niobium has revealed to be an appropriate material in the construction of superconducting cavities for particle accelerators. Due to the Skin effect, the electromagnetic waves (1-3 GHz) present in the cavity penetrate into the wall of the cavity and dissipate energy by Joule heating. This leads to numerous undesirable effects like loss in the Q factor of the cavity, loss in the beam definition and possibly quenching. It is therefore necessary to understand the heat transfer mechanism between the Nb wall and its surrounding He II coolant if optimum functioning of these cavities are to be maintained. In the experiment, which is based on a novel technique, we have considered not only surfaces of Nb that were prepared according to the standard procedure for the S.R.F. cavities, but also rough and chemically treated samples of different bulk purity content given by the RRR values. EXPERIMENTAL Experimental Cell Configuration : a performant test facility In an attempt to overcome problems, like determining the temperature of the solid near the interface, encountered in the standard technique for measuring the Kapitza resistance, we developed a new experimental configuration. Some of the advantages of this latter are : (a). The heat flux in the Nb is nearly one dimensional since the ratio of its diameter to thickness is relatively high. (b). The absence of thermometers on the sample implies an undisturbed uniform temperature field. (c). An easy and non-destructive sample mounting and dismounting procedure. The experimental set-up and cell are shown in figure 1. The experimental cell is composed of a stainless steel cylindrical support having an external diameter of 80mm and an internal diameter of 40mm. Two niobium plates, each 50mm in diameter and 2mm in thickness, are mounted on either ends of the support by means of two stainless steel flanges which are bolted together, The niobium plates are identical in that they were made from the same bulk material and their surfaces were prepared under the same conditions. An indium seal between the flange and the sample assures a superleak tight cell. This assembly forms a cavity which is filled with He II (called internal bath) through a stainless steel capillary tube of length l m and an internal diameter of 0.2ram. The experimental cell is immersed in a 4He cryostat ; helium surrounding the cell shall be referred to as the external bath. Manganin wire wound on a crosspiece made of epoxy constitutes an electrical heater resistor Rheaterl which controls the internal bath temperature. The latter is measured with an Allen Bradley 100W carbon resistor (AB 1). The temperature of the external bath, also measured with an Allen Bradley 100W carbon resistor (AB2), is controlled to within +0. l i n k with the aid of a heater resistance Rheater2. The Allen Bradley resistance thermometers have a high sensitivity in the experimental temperature range, that is, dR/dT -- 10 4 W.K -1 at 2K. 559
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ICEC16/ICMC Proceedings
Experimental Technique This is based on the fact that He II has a very small thermal diffusion time constant and that the Nb sample is at the same temperature as that of the external bath (To). The experimental procedure consists in measuring, under equilibrium conditions, the difference in temperatures between the intemal (Ti) and the external baths for different heat fluxes Q dissipated by Rheaterl. By symmetry, the heat flux through each Nb sample is Q', where Q ' = (Q - Qloss)/2. The difference in temperature AT = Ti- To and Q' are simply related by : (SAT/Q') = RG = RK + RNb
(1)
where RG is the global measured thermal resistance having two contributions 9RK the Kapitza resistance (or conductance hK) at the Nb-He II interface and the thermal impedance RNb = e/K, where e and K are the Nb sample thickness and thermal conductivity. The Kapitza resistance thus takes the form"
(2)
RK = (SAT/Q') - (e/K) = 1/hK
Now, Qloss is due to heat losses via the stainless steel support, the capillary tube and the electrical heater and thermometer wires. However, with the help of a numerical simulation these heat losses are quantified, thereby reducing the relative error in Q' (dQ'/Q') to < 0.2%. The thermal conductivity of each sample was determined by F. Koechlin [ 1] prior to measurements. Sample preparation Firstly, a series of measurements were done on Nb samples having rough surfaces (that is, as received from the supplier) with different bulk purity content. Sample 1 (RRR-~ 40), which was supplied by Wah Chang, was laminated from a sheet to a thickness of 2 mm. Sample 2 (RRR ,~ 180), which was supplied by Hereaus, was prepared as sample 1. Sample 3 (RRR --- 270) was machined from a rod supplied by Wah Chang. Secondly, to study surface effects on the Kapitza conductance, sample 2 was chemically polished according to the standard procedure for S.R.F. cavities. Finally, the influence of purification by titanium was investigated (samples 5 and 6 with RRR --- 370). RESULTS .
.
.
.
.
.
..------
J-nn io , Ua,
- -
--~
. . . . . .
11
r
Capillary ~ Experimental cell Internal bath Heating resistance
I
i i i i 1
] :~2.- -~', I [ ~-'~;--r.-',,i ,, i ! "~K,.._._..) u- / \r~ ._~~
(b)
i i i i
i i i 1
i i I i
i i i i i I i i i~
1
~ _-,,*---~---
eq |
R.e~malationthermomete _ / (AB2)
r !
~~L
T ~--~" ; .~J
Stainless steel flange :_... Niobium plate
=r--
:: - " ~ " . " : " ' : ' - ' - " - : " L ' T ~ T M
~
~ ! ~~-----~---
--
...~ .
He filling capillary Cylindrical support
l ' ~ ~ ' - ~.~:."-i'~"~i.?LLLd:: ~ - '.'?L~..: ' ' -..~-"'.'. '. -. .-. . . .
Manganin heater
~----~-~--~'~..,,..l.
Niobium plate
i
!,!
..
l ~-~
; -"Stainless steel flange
r
9
o
0,1
~
"~
E-..........]..............! I t
"----
J
i i .,i
+ II
9
O
sample 6,7 sample 5 sample 3 ........ sample 1 sample 4 sample 2
~ 1 7 61
1,5 i
,
! i i
.....
1,6
1,7
1,8
1,9
2
2
T(K)
I
Fig. 1 9Experimental set-up and cell
Fig. 2 9Kapitza Conductance of different Nb samples
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561
The experimental results of the Kapitza conductance between Nb and He II are shown in fig.2 for seven samples. The experiments were conducted at T < 2.1K and at saturated vapour pressure. The Kapitza conductance follows a T 3 behaviour for all samples as shown by the fits (solid lines), independent of their surface state. But, the magnitude of the conductance is between 25 to 50 times larger than the Khalatnikov's Acoustic Mismatch theory [2], shown by the solid line (a). The dotted line (b) represents the upper limit for hK as predicted by the phonon radiation model [3] for a perfect transmission across the interface. Effect of bulk purity. As seen in fig. 2 the bulk sample purity, indicated by the RRR value, does not have an apparent influence in the heat transfer process. For example, sample 3, which has a R R R value of 270, has a lower thermal conductance than sample 1, which has a RRR value of 40. This suggests that the discrepancy in the different sets of data arises mainly from the factors which influence the state of the solid surface, that is, its chemical and physical properties, and eventually the nature of the He close to the boundary. Influence of surface chemical polishing The effects of a chemical treatment were studied on samples 2 and 4 which are represented by open and full circles in figure 3. These two samples are identical, except that sample 4 was polished chemically as described in ref.[4]. Figure 3 shows that the heat transfer across a polished surface increases by a factor 2. Observations made in experiments by Mittag [5], whose results are also indicated in this figure (by crosses), show that not only heat transfer is enhanced, but a change in the temperature dependency may occur as well with polishing. The actuel heat transfer mechanism which is improved with polishing remains unknown. However, our results suggest a decrease in the reflection of the solid phonons at the boundary. Effects of Titanification Figure 4 shows the effect of surface impurities in heat transfer. Sample 5 (RRR -- 370) was coated with a layer of titanium. The Kapitza conductance of sample 5 (open diamonds) is close to that of rough surfaces. Now, approximately 3 mm from the surface of sample 5 was removed by chemical polishing (sample 6). We recall that surface oxides combine with Ti, a group 5 metal like Nb. Removing the Ti layer therefore
1,2 L .... 1
o
s
0,6
r'"
......
I''""
......
""l
....
_1
................................................................... ]........ ~ .............
i
!
i I - ........... i........~ !
~
i
i
.......i............. i.............. i. ~ . ~ " ~ i
i
.,X"
......
. . .........
"~
~
= 0
~ N
0,2
hd
1,4
f.t
6 0,4 r,.) "~
"
...... !iT............ i ..... -....... ;.~"~x--'~. - ' ' x 7 ' - ' i .............
' - ~ - X"- ~K- -~- -X- -.,~- XI~o~gh Nb s~mple (rcf.6) 0
....
1,5
i ....
1,6
!!
1,6
i ....
1,7
i ....
1,8
i ....
1,9
T(K)
i ....
2
00,6 ,6 4 0,4
0,2,2
!!!iiiii
1,5
i ....
2,1
00,8 ,8
2,2
Fig. 3 9Kapitza Conductance after chemical treatment
1,6
1,7
1,8
1,9
T(K)
2
2,1
2,2
Fig. 4 9Influence of purification by titanium
eliminates the presence of oxides on the surface of sample 6. The Kapitza conductance of sample 6 (solid diamonds) is a factor 2.2 larger than that of sample 5. Also, this conductance is about 25% larger than a sample which is only chemically polished. Another 45 mm (sample 7) was removed from the sample 6 surface and the Kapitza conductance remained unchanged in comparison with sample 6. This suggests once again the influence of surface boundary on the Kapitza conductance. Further, from a technological point of view, removal of a few mm of the surface after titanification suffices to improve the heat transfer.
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ICEC16/ICMC Proceedings
CONCLUSION The experiment shows that heat is evacuated more easily across a Nb-He II interface than expected on the basis of the Khalatnikov theory. The results suggest that this is due to the nature of the boundary rather than bulk properties of each medium. The Kapitza conductance for rough surfaces is of the order of O.057xT3 (W.cm-2K-1). The conductance is a factor --2 larger for surfaces that underwent a chemical treatement. Finally, for samples that were treated for impurities by a titanification process, the Kapitza conductance is of the order of O.15xT3 (W.cm-2K-1), that is, heat is removed at a rate of 0.86W.cm-2.K -1 at T = 1.8K, the optimum temperature of a S.R.F. cavity. REFERENCES
,
,
4. 5.
Koechlin, F., Proceeding of the International Conference on Superconductors, ICMAS-90, Grenoble, France (1990) Khalatnikov, I. M." Introduction to Superfluidity, W. A. Benjamin, Inc., New York, New York (1965) Snyder, N. S., Heat transport through helium II : Kapitza conductance Cryogenics (1970) 10 89-95 Boucheffa, A., Th~se de l'UniversitE Paris 6 (1994) Mittag, K., Kapitza conductance and thermal conductivity of copper niobium and aluminium in the range from 1.3K to 2.1K. Cryogenics (1973) 10 94-99
Thermal Behaviour of Electrical Multilayer Insulation Permeable to Superfiuid Helium B. Baudouy t , A. Boucheffa*, M.X. Francois* and C. Meuris CEA/Saclay, DAPNIA/STCM, F-91191 Gif-sur-Yvette Cedex *CNRS/LIMSI, B~tt. 502 ter, Campus universitaire, F-91405 Orsay
Electrical multilayer insulations made of Kapton | tapes and prepreg or adhesive Kapton ~ tapes, used in dipole magnets, may offer a complicated arrangement of thin helium channels which cannot be easily predicted and modelled. Several insulation systems have been tested in order to characterize their helium channels pattern. Heat transfer data analysis shows clearly the contribution of supeffluid channels inside the insulation. Appearance of the vortex-free regime for very small temperature differences (10 .5 to 10.3 K) and of the Gorter-Mellink regime for higher temperature differences allows to establish the mean value of the channels diameter. We present in this paper the thermal behaviour of several combinations of insulating materials with different geometrical arrangements and porosities.
INTRODUCTION The research and development program for the Large Hadron Collider dipoles developed at CERN includes stability studies which are carried out in collaboration between CERN and CEA/Saclay. For NbTi magnets cooled by superfluid helium the most severe heat barrier comes from the electrical insulation of the cables. This paper reports a work which is part of the thermal study program. It deals with the intrinsic qualification of different insulation systems. Their global thermal performance in the surroundings of the winding is studied with a different experimemal model [ 1]. Classically, an insulation is a composite made up of a first wrap, a tape wound around the cable for electrical insulation, and a second outer wrap protecting mechanically the inner wrap, creating helium channels and gluing to the next conductor to keep the coil in shape. In the tests described, wraps are reproduced as plane layers. One (samples B22 and B23) or two (sample B25) sublayers of 11 mm wide Kapton | tapes with an overlap of 50 % are used for the first layer. For the second layer, adhesive Kapton ~ tapes, 12 mm wide with a spacing of 2 mm (B23 and B25) or 4 mm (B22), are employed.
THERMAL QUALIFICATION OF INSULATION Two 50.3 cm 2 wide insulation samples are clamped between two isothermal baths of He iI, each of them being able to reach different temperature. A detailed description of the experimental set-up is given in [2] and summarized in figure 1. Temperature measurement is made after attaining a stationary temperature for both baths, the outer one being regulated and held constant for the whole test over the range of power dissipation values in the inner bath. The difference in temperature between the two baths is plotted as a function of the generated power. This curve, of which an example is given in figure 2a, characterizes the overall thermal resistance between the heated bath and the cold bath. The measured resistance includes the Kapitza resistance, the resistance of the insulating material and the resistance of the helium paths through the insulation samples. Tests have been performed at Saclay in a pressurized He II cryostat. Preliminary results reported in [2] have shown that heat transfer is influenced by He II heat transfer even for small volume of helium inside insulation. At low heat flux, heat transfer is purely governed by He II. Fits in this range of heat flux Doctoral fellow CEA-Jeumont lndustrie-CERN 563
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ICEC16/ICMC Proceedings
have allowed to determine an equivalent geometrical factor A/L1/3 which characterizes a permeability of the medium related to He II heat transfer in a net of narrow channels (equivalent cross-sectional area A, length L) opening on either side the insulation. For higher heat flux, a model which considers insulation material and He II thermal paths in parallel agrees with measurement within 10 %.
Figure 1 Schematic diagram of insulation and experimental setting. Dimensions are in mm
EXTENSION OF MEASUREMENTS TO VERY LOW TEMPERATURE DIFFERENCES In another test facility at LIMSI, with a very precise temperature controlled saturated He II bath, using a resistance a.c. bridge equipped with a lock-in null detector, it has been possible to measure temperature difference as small as 10 laK across the same insulation samples. Figure 2b shows an typical example of the inner bath temperature rise AT, for different temperatures of the outer cold bath Tb.
Figure 2 Temperature difference between the two He II baths as a ftmction of the power dissipated m the inner bath for various outer bath temperatures. Sample B22. (a) : range [1 mK, AT~ = T~ -Tb]. (b) : small AT
MEASUREMENT ANALYSIS Heat is transported in He II according to the movement of the normal fluid from the hot source towards the cold source. The motion of this viscous fluid undergoes a resistance whose magnitude varies according to the associated Reynolds number and the possible presence of vortices in the superfluid component.
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Very often the heat flux densities q = Q/A are such that the double convection velocities v = v, v~ exceed the thresholds for the formation of vortices. The mutual friction generated becomes thus dominant and leads to a transport law known as the Gorter-Mellink regime [3,4], IqlZTq--f(T)VT, where f(T)-
(9~3s4Ta)/(A(T)9,) with the notations used by the authors stated, that leads to"
Q3_ ~
~f(T)dT L
(1)
In media of small dimensions, the critical velocities for the formation of vortices become sufficiently high such that the corresponding critical flux densities are non-negligible [5]. This gives a small range in which heat is transferred by a normal fluid without any interaction with the superfiuid component. The Landau two-fluid model without interaction [6], leads to the transfer law V p - p s V T , with ~'p - -(12g./d2)'~ for narrow slits of thickness d in the Poiseuille regime, that is" Q - (Ad2)(ps)ET 12~t~ AT
(2)
The analysis of the experimental results, of which an example is reproduced in figure 3, indicates : 9 the importance of the He II contribution to heat transfer by an estimation of the purely conductive part through the solid structure, Q~ol, from measurements of the conductivity and of the Kapitza resistance, which were done in the same experiment cell on samples of plain Kapton | foil 9 the possible existence of a laminar Landau regime 9 in the affirmative case, the determination from equation (2) of the geometric factor Ad2/L - X~' Aidi2/L, equivalent to the totality of n micro-channels 9 the verification of the hypothesis by the temperature independence of this geometric factor 9 an estimated value of the critical heat flow Qr and the corresponding superfluid velocity v~ 9 the Gorter-Mellink regime by a research of a zone (in which Q~o~is ever lower than 0.1 Q) having a Q3 dependency and, within the precision of fiT), a value of the geometric factor A/L v3 - E~' A~/L~ v3, after equation (1), and finally, the verification of its non-dependence with temperature.
Figure 3 Different heat transfer regimes. Sample B22. The solid curves are the best fits (including measurement errors) to the data using equations (1) and (2). The dashed line is the estimated purely conductive part Those different results, extracted from the AT(Q) curve of sample B22, are reported in table 1. We note that the geometrical coefficients corresponding to the two regimes stated above are almost constant in a large temperature range. It is possible to deduce values of A and d, that is A = 5.74 mm 2 and d = 2 2 ~tm,
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ICEC16/ICMC Proceedings
by making the hypothesis that the He II slits are of length 5.5 mm corresponding to the overlapping of the tape of first insulation layer, and by supposing identical dimensions for all the parallel slits. The critical flux densities QdA vary from 0.32 W/cm z to 0.68 W/cm 2 when the temperature of the bath varies from 1.7 K to 2.0 K. The critical velocities (superfluid values) vary from 1.0 cm/s to 3.1 cm/s and the associated Reynolds numbers from 156 to 100. We thus verify that the normal fluid flow is laminar in the Landau regime. Table 1 Equivalent geometric factors of the insulation sample B22 Temperature Ad2/L 10"14m 3 A/L 1/3 10"5m"~/3
1.7 K 53 + 51 3.29
1.8 K 53 + 7 3.20
1.9 K 49 _+6 3.25
2.0 K 41 _+9 3.17
2.05 K
3.25 1 The precision includes the precision in the measurementsand the precision in the fits of Landau and Gorter-Mellink zones. The preceding results obtained for different insulations are summarized in table 2. For sample. B25, the slits are very narrow, the wall friction of the normal fluid induces a non-negligible temperature gradient which does not allow a good identification of the Gorter-Mellink regime. We observe that the results of B22 and B23, for which L is a priori identical, brings to front a factor 2.6 on A which is coherent with the ratio 2 on the spacing of the second layer tape and the fact that the thickness d of the slits also decreases when the spacing decreases. The doubling of the first layer (B25) can be understood by a strong reduction of the effective permeability, which we must attribute not only to the increase of the channel lengths but equally to a decrease of d.
Table 2 Characteristic geometric parameters of the He II slits of different insulation systems Sample first layer Kapton| spacing of the number tape (50 % overlap) second layer tape 2 B22 150 HN (38 lam) 4 mm B23 150 HN (38 lam) 2 mm B25 2 x 100 HN (25 lam) 2 mm 2 Kapton| 270 LCI tape (adhesive, 68 lam)
Ad2/L
10"14m3 49 7.1 0.18
A/L1/3
10"-~m~/3 3.25 1.19
(A/L 1/3)/mtot
10-3m"1/3 3.23 1.17
CONCLUSION The results and the method of analysis presented allow a detailed qualification of insulations by determining characteristic geometrical magnitudes of thermal properties and, in particular, that of parameter d, the thickness of slits which can only be attained experimentally.
ACKNOWLEDGEMENTS The authors are grateful to Dr D. Leroy and Dr B. Szeless from CERN for helpful discussions. Special thanks to Mrs A.M. Puech and Mr Gaubert for the sample preparation and measurements.
REFERENCES 1 Burnod L. et al., Thermal modelling of the LHC dipoles functioning in superfluid helium, Proc. EPAC C one (1994) 2 Baudouy B. et al., Steady-state heat transfer in He II through porous superconducting cable insulation, Proc. CEC-ICMC Conf. (1995) 3 Vinen W.F., Mutual friction in a heated current in liquid helium II, Proc. Roy. Soc. A240-A243 (1957) 4 Gorter C.J. and Mellink J.H., On the irreversible process in liquid helium II, Physica 15 (1949) 5 Arp V., Heat transport through helium II, Cryogenics 10 (1970) 6 Landau L.D. and Lifshits E.M., Statistical Physics, Pergamon, Oxford, UK (1958)
Pressure Gradient Caused
by
Quantized
Vortex
in
Superfluid
Helium
Minoru Yamaguchi, Yoshiko Fujii, Masaki Nakamura, Toshinobu Shigematsu and Toyoichiro Shigi Dept.of Applied Physics, Okayama Univ.of Science Ridai-cho,
Okayama
700,
Japan.
The temperature difference and the pressure difference through He II in a capillary glass tube have been measured very precisely.
The quantized vortex
line density was calculated from the temperature difference data using the numerical scaling coefficients.The pressure difference data give the information of interaction between the vortex line and the tube wall.
The pressure
dissipation caused by the vortex line was only one tenth of mutual friction between the vortex line and the normalfluid.
In the terms of eddy viscosity for
superfluid velocity field, we have calculated the coefficient of the interaction between the vortex line and the superfluid. INTRODUCTION A great deal of experimental and theoretical studies on superfluid turbulence have been performed [1],[2],[3],[4],[5], however, the influence of the vortex line on pressure dissipation is still in question. To study this problem, we have measured the temperature andpressure difference across the glass capillary tube in which the thermal counterflow was built up. EXPERIMENT Fig. l shows schematically the arrangement of the experimental cell to carry out the temperature and pressure difference measurements. Heat supplied in lower part of the cell produces thermal counterflow in a capillary glass tube.
During a series of measurement,
the temperature of upper part of the cell was always maintained constant within
I x 10SK.
The temperature difference across the tube was measured with the accuracy of 1 p K using a thermocouple (Au[0.03at.%Fe]-NboTi) connected to a SQUID detector.
Tow membrane type pressure
transducers were used to measure the pressure. The 8 ~ m-thick phospor bronze membrane and the transducer body form a electric capacitor.
Fig. 1 Schematic arrangement of this experiment. 567
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ICEC16/ICMC Proceedings
The value of pressure is obtained from the resonant frequency(f) of a tank circuit connected to a tunnel diode BD-5.
The stability of
able to be detected.
NO. 1 2 3 4 5
was about 106 , so that the pressure change of
1 • 10-2 Pa was
The size of the tubes used in this experiment is listed in Table 1.
Table tube
f
1
Table
d (x 10-2cm) 1. 3 0 I. 2 6 i. 8 0 1. O0 2. 6 6
1 (cm) 9. 6 9. 6 9. 5 9. 7 9. 5
T (K) 1. 3 i. 4 i. 5 1. 6 1. 7 1.8 1.9
2
(Ref.[7])
CL 0. 080 O. I 0 0 O. i18 O. 1 3 3 O. 1 5 4 0.174 0.195
(I ,I--CLI1) 0.710 0. 710 0. 713 0.719 0.731 0. 742 O. 7 5 5
RESULTS AND DISCUSSIONS The temperature differences (AT) as a function of the heat flux (W) for tube NO.2
are shown in Fig.2.
If the heat flux is less than W~, AT is expressed by a following relation,
128r/.1
AT
w
nd
(])
7'
where r/,, is the viscosity coefficient of the liquid helium, 1 is the tube length, d is the inner diameter of the tube, p is the density of liquid helium, S is the entropy density and T is the temperature of the liquid helium respectively.
Hence, in this state the normal component of liquid helium is considered
to be the laminar flow. In the turbulent state, the temperature gradient across the tube V T is represented as follows[6],
VT-
VI"t = ~ P.fi
where
,
(2)
t~[,, = p.~xct(Ill - c t I , ) L V
VT~ is caused by the contribution of the normalfluid in the laminar state, F,, is the mutual
friction force between the vortex line and normalfluid, p.~ is the superfluid density, K is the quantum of circulation, a is the interaction coefficient, L is the vortex line length density, V is the relative velocity of two fluids. (I It - c t I~) is the parameter of the dynamical scaling method. L was calculated from f l L v2
VT
data using the numerical value for (Ill-crib), [7].
Fig.4 shows the relations of V and
for tube NO. 1, where fl is represented as follows [6],
13 = (
) In(aLl/2 )
( where,
a is the radius of a vortex filament ).
In the well developed turbulent state f l L ~/2 can be represented as follows, pL !/2 - -
CL2
V
The coefficient CL2 obtained
(3) from the experimental data is shown
The value of ci. 2 is independent on the tube diameter. vortex line length.
for tube NO. 1,2,3,4,5
in Fig.5.
Hence, we should consider that L is the net
ICEC16/ICMC ICEC 16/ICMC Proceedings Proceedings tube N0.2
0
I
I
569 569
I
~..*. .*
A
100
0
W
200
( p watt)
Fig2 Temperature Werence A7' aB a
function of the heat flux W for tube N0.Z.
n 1
0.2 N A.
I
0
v
t
2Q
10
l c m - s-€1
Fig4 The vortex Pine density
3
for tube NO.1.
an a function of Y
t
n
0.1
"
1.2
1.6
T m
1
3
(K)
F i g 5 Coefficient cLZ as a function
c
of the temperature for the tube NO.1,2,3,4,5.
v
(cm-
h
s-1)
Fig.6 The i n d u d premure gradient 8s a function
.
Icr
.
p PPr
of Y for tube N0.1.
.
i
o
a
e
o
,
Q.001
E .6 20 T IK3 Fig.8 T h e parameter & as a function of f .2
the temperaturefor tube NO.1,2.3,4.6.
as
CX103
* $">
Fig.7 V f ; aa a function of PI4 for tube N0.1.
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The pressure differences (AP) are shown in Fig.3 for tube NO.2. In the turbulent state, the pressure gradient across the tube may be represented as follows,
(4)
V P - Vl"L oc C~psM~V
where VPL is the contribution of the normal fluid in the laminar state and c~z is any coefficient of the interaction between the vortex line and superfluid. For the experimental data., the coefficient c a does In this experiment, flVP r increases in proportion to V 2 as shown in
not remain in a constant value. Fig.6(where VPr = V I " - VP s ).
Then, we assume that the pressure gradient is given as follows, 32r/, VI~. = - ( d 2 )zV,, where y is a characteritic parameter.
(5)
The form of r/, may be assumed as follows[2],
r/~ = p~x~(d)L ';2
(6)
where ~ is a fitting parameter.
Then,
by means of (6), we have the following results.
VP r = -( 16ps2tc~Y )vL';2
(7)
pa
and when we substitute eq.(3) for V of eq.(7) [V/r [ = (16p.,2tc~y fl) pd )'( L CL2 VPr as a function of pL is shown in Fig.7. viscosity derived from VPr combined parameter ~y is Then
16p,~:
(pdLi, )
L~/2=2000 (cm-~),
(8) The combined parameter ~:y
data is shown in Fig.8. 3.5 • 103.
In this ~y,
For the representative resulut (in if
1.7K), the
we take Y' =0.5, then we get ~:=7.0 • 10-3 .
may be compared to a (I II --CLI;)we can calculate as follows,
of superfluid eddy
When
y =0.5,
~ =7.0 • 10-3
and
16p~ (pdLiJ2)=0.002.
This value is about one tenth of the value of a (I)l -cs REFERENCES [ 1] Vinen,W.F.,Proc.R.Soc.London ser.A242(1957). [2] Brewer,D.F.and Edwards,D.O.,Proc.R.Soc.London,ser.A251 (1959)247;Phios.Mag.6(1961)775;6(1961) 1173;7(1962)721. [3] Childers,R.K.and Tough,J.T., Phys.Rev.B 13(1976) 1040-1055. [4] Martin,K.P.and Tough,J.T., Phys.Rev.B27(1983)2788-2799. [5] Tough,J.W.,Superfluid Turbulence,Progress in Low Temp.Phys.,edited by D.F.Brewer(North-HoUand), (1982)8,Chap.3. [6] Donnelly,R.J., Quantized Vortices in Helium II, Chap.7 (p.215-254), (Cambridge Univ. Press.) [7] Schwarz,K.W., Phys.Rev.B38(1988),2398.
Heat and Mass Transfer between Two Saturated He II Baths X. Huang, J. Panek, and S.W. Van Sciver National High Magnetic Field Laboratory, 1800 E Paul Dirac Dr., Tallahassee, FL 32306, USA The present paper examines the heat and mass transfer processes between two saturated He II baths. The two baths, formed from two stainless steel cans of 50 mm ID and 660 mm length, are connected at the top and the bottom with two 100 mm long tubes. The bottom tube, with a 5 mm ID, is filled with liquid He II; the top tube, with a 1.3 mm ID, is filled with saturated He vapor. Heat transfer between the two baths is then governed by the counterflow process in the bottom tube and vapor mass transfer process in the top tube. Steady state and dynamic models, based on energy and momentum equations, are presented and agree very well with experimental results. Our study also demonstrates that the mass transfer process is a far more efficient heat transport mechanism than the counterflow process in vapor/He II two-phase systems.
INTRODUCTION A new cooling scheme using two-phase vapor/He II has been proposed for the superconducting dipole magnet strings in the Large Hadron Collider (LHC) [ 1]. The objective is to absorb the heat generated at 1.9 K in the LHC with minimum temperature rise by utilizing the latent heat of liquid He II. Such a cooling scheme, however, suggests two-phase flow and heat transfer in saturated He II over the whole range of vapor qualities. Unfortunately, despite preliminary experimental tests performed at CERN demonstrated the feasibility of such a cooling scheme[2], little has been know about the basic heat transport mechanism in two-phase vapor/He II system. To better understand the basic heat transport mechanism in a two-phase He II system, we conducted analytical and experimental studies on the heat and mass transfer between two saturated He II baths. Because the counterflow heat transfer in the liquid and mass transfer in the vapor are decoupled in this problem, we were able to evaluate the two transport processes individually and thus identify the dominate mechanism in the system. In the present paper, we present the analytical solution to the steady state process as well as the numerical modeling of the transient process of the system. Experimental results are shown to be consistent with the predictions based on the theoretical models. EXPERIMENTAL APPARATUS A schematic of the experimental apparatus is shown in Figure 1. Two stainless steel cans, 5.0 cm in ID and 66 cm in length, are connected at the top and the bottom with two 10 cm long ss tubes. The top tube has a 0.13 cm ID and the bottom tube has a 0.5 cm ID. The two cans are installed inside a vacuum can which isolates the two cans from the He I bath. Liquid helium is supplied to the cans through a JT expansion valve from the He I bath. A 56 cm long coaxial capacitance type liquid level meter is used to monitor the He II level inside the cans. Two germanium resistance thermometers, calibrated in situ against the helium vapor pressure, are placed inside each can. Analysis of the calibration suggests the error associated with the temperature measurements is less than 1.0 mK. A 170 ~ resistive heater, located at the bottom of the left side bath (the warm bath), is used to supply either steady state or transient heating to the warm bath. The right side bath (the cold bath) has a 25.4 mm ID pump-out line at the top connected to a high capacity vacuum pump through a control valve. Differential pressure across the two He II baths is measured with a variable reluctance differential pressure transducer. The experiment procedure begins with filling both baths with liquid helium to about half the height of the cans. With the JT fill valve closed, the vapor pressure inside the He II baths is then regulated to a prescribed value. In a steady state experiment, a constant power is supplied to the warm bath resulting in temperature and pressure increases in the warm bath. The pressure increase in the warm bath then forces liquid to flow to the cold bath until the difference in the hydrostatic head between the two baths offsets the vapor pressure difference between them. Data are taken when the temperatures in both baths reach steady state. For the transient experiments, because of difficulties in maintaining a constant vapor pressure in the cold bath during a transient heating, the control valve is 571
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closed and constant volume is maintained. The dynamics of this closed volume system is then studied by depositing a square heat pulse of 1.0 s duration to the warm bath and recording the temperature, pressure, and liquid level traces. ANALYTICAL MODELING Steady State Process Because of the extremely high effective thermal conductivity of He II, it is assumed that the temperature of each He II bath is uniform and equals to the saturated temperature at the corresponding vapor pressure. The steady state heat transfer between the two baths is governed by the counterflow in the bottom tube and the mass transfer in the top tube. For turbulent vapor flow in the top tube, the equations are given as
Qc =
f l (T) dT
/'
(1)
and (Ps(Th)- ps (Tc))pgdg 1"25/ 4/7 0.158L//0.25
Q - Qc = Aghfg
(2)
where Q is the total heat deposited in the warm bath, Qc is the heat conducted away by the counterflow process through the bottom tube, ps(T) is the saturated vapor pressure at temperature T, Af is the cross sectional area of the bottom tube, Ag is the cross sectional area of the top tube, L is the length of the tube, hfg is the. latent . heat .of evaporation, pg is the average vapor density, dg is the top tube inner diameter, and # is the vlscosxty of the vapor. Transient Process During a transient process, the changing temperature and pressure in each bath result in liquid level changes in each bath. Consequently, the energy transfer between the two baths not only involves transient Gorter-Mellink conduction and vapor mass transfer but also mass transfer by the liquid in the bottom tube. The energy balance for each bath during a transient process can be written as
]
~- pfVfiefi + P g i ( V - Vfi)egi = Qi -+ riaghg _+rhfhf _+Qc
(3)
where the plus sign is for the cold bath and the minus sign is for the warm bath. The subscripts f and g designate liquid and vapor, respectively. V is the volume of the can, Vfi is liquid helium volume in the bath, e is the specific internal energy. Qi is the external heat deposited in the bath, Qc is the energy transfer by the transient Gorter-Mellink conduction. The vapor flow in the top tube and the liquid flow in the bottom tube are governed by the one dimensional continuity and momentum equations cgp t- cg(pu) _ 0
&
anu
(4)
0x
~u
2Cf PU2 _ K p u 2 _ ~
P'&- +PU~x = -
d
(5)
where Cf is the frictional factor for the fluid, d is the tube diameter, and K is the coefficient for the entrance effect. The transient counterflow heat transfer for the warm bath is
~]1/3 (6)
Qc = A f [ f l ( T ) ~
x=0 For the cold bath, the derivative should be evaluated at x=L. To evaluate Eq.(6), the temperature profile along the bottom tube is needed. This temperature profile can be obtained by solving the He II energy equation,
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o,-tr rhfCp onrr Cp oqt _--4
A f p f c3x
3 i rl(T) _~]1/3 =0
oax
(7)
where9 C P is the specific heat of liquid He II. The final equation needed is the mass conservation equataon for each bath,
]
~- pfV fi + Pgi (V - V fi ) __.rhg 4- rhf = 0
(8)
where the plus sign is for the cold bath and the minus sign is for the warm bath. The initial conditions for these equations are
Tc=Th=Tinit, rilg -- rilf = 0, Wfi -- Vinit , at t=0. where Tinit is the initial He II bath temperature and Vinit is the initial He II volume. RESULTS Figure 2 displays the steady state relationship between the heat input and the temperature difference between the two baths at 1.9 K along with calculations based on Equations (1) and (2). Two important conclusions can be drawn from this figure. First, although the upper tube has a flow area about 15 times smaller than that of the bottom tube, the heat transferred by the mass-transfer process through the top tube is still greater than that by the counterflow process over majority of the heat input range. This clearly suggests that the mass transfer process is a far more efficient heat transfer mechanism than the counterflow process. Second, the experimental data agree very well with the model prediction especially at power input below 0.9 W. Figure 3 plots the temperature traces for the warm and cold baths after a 10 W square heat pulse of 1.0 second duration was sent to the warm bath. The initial liquid helium level is 37.5 cm and the initial bath temperature is 1.790 K. The curves are numerical solutions to Equations (3) through (8) with helium propertiesevaluated by HEPAK. During the initial heating, the temperature in the warm bath increases rapidly and reaches the maximum at t=l.0 s when the heater is switched off. Subsequently, the warm bath starts to cool down while the cold bath continues to warm up until both baths reach the same temperature at t = 15 s. Figure 4 shows the measured liquid level in the warm bath as a function of time along with the numerical prediction. When the transient heating is initiated, the rising pressure inside the warm bath forces the liquid to flow to the cold bath and thus the liquid level drops. At about t=5 s, the difference in the hydrostatic head equals to the vapor pressure difference between the two baths and the liquid flow reverses. When the two baths finally reach the same temperature at t -- 15 s, the liquid flow also stops and the liquid level in the warm bath almost returns to its initial value (the amount of liquid evaporated by the transient heating is negligible). CONCLUSIONS We have demonstrated the behavior of two-phase He II/vapor heat and mass transfer in an idealized experimental configuration. The steady state result clearly suggests that the vapor mass transfer is the dominate heat transfer mechanism in two-phase He 13Jvapor systems. The dynamic model presented in the paper agrees very well with the experimental results. Our work continues with a study of the fully coupled horizontal two-phase He II flows. ACKNOWLEDGMENT The authors wish to thank Reda Daher and Vincent Cochran for their valuable technical support. This work is supported by the U.S. Department of Energy - Division of High Energy Physics under grant DOE-FG02-96ER40952 REFERENCES 1. 2.
P. Lebrun, Superfluid helium cryogenics for the Large Hadron Collider Project at CERN, Crvozenics(1994) 34 1-8 A. I3ezaguet, et al., The superfluid helium model cryoloop for the CERN Large Hadron Collider, Adv. Cryo. Engr. (1994) 39, pp 649
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Figure 1. Schematic of experimental apparatus
Figure 2. Steady state heat input versus bath temperature difference at 1.90 K
Figure 3. Temperature traces of the warm and cold baths after a square heat pulse. power=-10.0 W, duration=l.0 s.
Figure 4. Liquid level in the warm bath after a square heat pulse to the warm bath. power=-10.0 W, duration=l.0 s.
Measurement of Characteristic Time of Quantized Vortex Development Using a Thermal Shock Wave
Takeshi Shimazaki and Masahide Murakami Institute of Engineering Mechanics, University of Tsukuba Tennoudai 1-1-1, Tsukuba, Ibaraki 305, Japan
The characteristic time of quantized vortex development is experimentally measured by using a propagating thermal shock wave. A thermal shock wave generated by a pulsed heating from a planar heater is deformed through the interaction with quantized vortices. The deformation is measured by means of a superconductive temperature sensor. It is found tha, t the characteristic time is inversely proportional to the square of applied heat flux. It is also found that appreciable effect appears on transient heat transfer if the vortex line density exceeds approximately 105cm/cm a.
INTRODUCTION He II is regarded as an excellent coolant for such as superconducting magnets. However, there are still several unsolved open questions concerning highly transient heat transport in it. It is indispensable to make those questions clear in order to understand the heat transport properties for the practical applications of He II. One of the most important and complicated questions is how the quantized vortex lines behave. When the relative velocity between the superfluid and normal fluid components exceeds certain critical value, quantized vortices are induced and the dissipative effect due to the interaction with quantized vortices get to be appreciable on heat transport phenomena. For steady or quasi-steady cases, the approach introduced by Gorter and Mellink[1] is usually used to take into account the effect of quantized vortices on heat transfer and is believed to give fair results both qualitatively and quantitatively. It is, however, known that the approach loses the validity for highly transient cases. In the cases that the time scale of thermo-fluid dynamic phenomena becomes comparable to or shorter than that of the characteristic time of evolution of quantized vortex lines in such as highly transient cases, their development and decay should be taken into account. The development and decay of quantized vortex lines were first successfully formulated by means of phenomenogical approach by Vinen[2,3,4]. The Vinen vortex line density equation has been widely used in many investigations[5,6]. In this study the characteristic time of quantized vortex development, defined from the point of view of a transient heat transfer, is derived by analyzing the thermal shock wave profiles measured with a superconductive temperature sensor. The result is also compared with the numerical result based on the vortex line density equation.
EXPERIMENTAL SETUP AND PROCEDURE Whole measurements are carried out in He II under the saturated vapor pressure condition. Figure 1 shows the main experimental apparatus itnmersed in He II. It consists of three ma.in parts, a superconductive temperature sensor, a. planar Ni/Cr thin film heater and a cylindrical 575
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side wall or a shock tube[7]. The distance from the hater to the sensor can be varied between 0.1 to 150 m m with 0.1 turn step. The sensing element of the sensor is superconductive thin metal film consisting of gold and tin fabricated on to a side wall of quartz fiber (1.5 m m in length and 40 #m in diameter). Constant current is applied to the element and the temperature variation is measured as an variation of voltage drop across the element due to the superconductive transition. Trapezoidal current pulse is applied to the Ni/Cr thin film heater which is 27 m m x 27 mrn and 400 ~ in thickness to generate a thermal shock wave. The cylindrical side wall, 150 rnm in length and 25 turn in inner diameter, maintains the one dimensional character of a traveling thermal shock wave inside the wall. The characteristic time of quantized vortex development tv~ is obtained by analyzing thermal shock wave deformation due to the interaction with quantized vortices. If the quantized vortex line density is not sufficiently high, a thermal shock wave profile changes only as a result of hydrodynamic nonlinear effect, that is to say the formation of a shock wave and the deformation into a right-angled triangular form. The applied heat from a heater is wholly transported by a thermal shock wave[7]. On the other hand, if the vortex line density becomes high enough, dense quantized vortices bring about dissipative effects on the heat transport, and a partial declination appears in the late portion of the plateau of a thermal pulse as illustrated in Figure 2. The time interval, t~, from the wave front to the point at which the declination of plateau becomes appreciable is defined as the characteristic time of quantized vortex development from the point of view of a transient heat transfer. The point of onset of declination is practically defined as the instant at which the deviation exceeds 3 cr of the data fluctuation.
RESULTS AND DISCUSSION Figure 3 shows typical measured thermal shock wave profiles for four cases of the heat flux %. It is seen that the point of declination onset gets closer to the wave front and the deformation becomes larger as qp increases. Figure 4 shows the measured t ~ as a function of % in double logarithmic plot for two cases of temperatures, 1.70 K and 1.90 K. The solid and broken lines, of which inclinations are -1.9 and -2.0, respectively, represent the linear regression results. It may be concluded that tv~ can be approximately related with qp in the following equation.
t~-
C ( T s ) q ; 2,
(1)
where C ( T s ) is the temperature dependent coefficient given experimentally. The results are also compared with the numerical result based on the Vinen vortex line density equation given by
dL L3/2 d--t- = a Iv,~[ - /3L2 + 7
Iv. l 5/2 ,
(2)
where v~ is the relative velocity between the superfluid and normal fluid components, L is the vortex line density defined as the total vortex line length per unit volume, a and /3 are the growth and decay coefficients given by Vinen, 7 is also given by Vinen[4]. The third term of the right hand side is the source term which is usually neglected. However, Kanari et al. [8] reported that the term is indispensable especially in the case of strong heating. Figure 5 shows the time evolution of quantized vortex line density obtained numerically by solving the equation without neglecting the third term for three cases of the applied heat fluxes, where v~, is given as a function of applied heat flux %. It can be seen that the final density reaches higher value and the development becomes faster for larger %. It is easily seen that the time duration which is required for the vortex density to reach certain value L is a function of heat flux. The time duration in which L reaches a number of values, from 103 to 107cm/cm 3, is plotted in Figure 6, where experimental data. a,re also plotted for comparison. It is seen that the inclina,tion of the regression result va.ries fl'om - 5 / 2 for the vortex line density of 103cm/cm :3
ICEC16/ICMC Proceedings to - 3 / 2 for 107cm/cm 3. The experimental data are found to reasonably follow the line of L = 5 x 1 0 4 c m / c m 3 of which declination is - 2 . It can be concluded from the figure that if the vortex density exceeds approximately lOScm/cm 3, transient heat transfer will be appreciably affected by quantized vortex lines.
CONCLUSIONS The characteristic time of quantized vortex line development is experimentally measured frolla the deformation of thermal shock wave profiles. It is found that the characteristic time t~, can be related with the applied heat flux qp as t~, - C(TB)q~ 2. It is also found that when the vortex line density exceeds the order of lOScm/cm 3, the transient heat transfer get to be affected by quantized vortex lines.
ACKNOWLEDGEMENTS This research was partly supported by JSPS (Japan Society for the Promotion of Science)
REFERENCES
1 Gorter, C. J. and Mellink, J. H. , On the irreversible process in liquid helium II. Physica (1949) 15 285-305 2 Vinen. W. F. , Mutual friction in heat current in liquid helium II. I. Experiments on steady heat currents. Proc. R. Soc. London A (1957) 240 114-127 3 Vinen. W. F. , Mutual friction in heat current in liquid helium II. II. Experiments on transient effects. Proc. R. Soc. London A (1957) 240 128-143 4 Vinen. W. F. , Mutual friction in heat current in liquid helium II. III. Theory of the mutual friction, Proc. R. Soc. London A (1957) 242 493-515 5 Murakami, M. and Iwashita, K. , Numerical computation of a thermal shock wave in He II. Comp. & Fluids (1991) 19 443-451 Fizdon, W., Schwerdtner, M. v., Stamm, G. and Poppe, W., Temperature overshoot due to quantum turbulence during the evolution of moderate heat pulse in He II. J. Fluid Mech. (1990) 212 663-684 Shimazaki, T. , Mura.kami, M. and Iida, T. , Second sound wave heat transfer, thermal boundary formation and boiling: highly transient heat transport phenomena in He II. Cryogenics (1996) 35 645-652 Kanari, T. and Murakami, M. , Numerical investigation of evolution of vortex line density in the case of transient heating, to be presented at ICEC 16 (1996) 9 Nemirovskii, S. K. and Tsoi, A. N . , Transient thermal and hydrodynamic processes in superfluid helium. Cryogenics 29 985-994
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Figure 2 Schematical drawing of the definition of Figure 1 Schematical drawing of the main the characteristic time of quantized experimental apparatus immersed in He II. vortex line development tve.
Figure 3 Superposed thermal shock wave profiles generated by the trapezoidal heat pulses of various heat flux value qp. The temperature T B is 1.90 K. The heating time tH is fixed at 500 Its. The distance from a heater to a sensor z is fixed at 10 mm.
Figure 5 Time evolution of quantized vortex line density obtained numerically by solving the vortex line density equation without neglecting the third tenn. The heat flux is supposed to be applied during the calculation. The initial VLD is assumed ,.) to be 10-cm/cm 3 for every case. (The calculation is done by Kanari et al. [8])
Figure 4 Measured characteristic time of quantized vortex line development tve as a function of applied heat flux qp for two temperatures. Linear regression results are also shown in the figure.
Figure 6 Comparison of the characteristic time tve and the time duration required for the quantized vortex line density L to reach certain values.
Cryogenic engineering
Heat transfer
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Non-dimensional Correlation for Boiling Heat Transfer from Sintered Porous Layer Surface
Rongshun Wang, Anzhong Gu, Zhen Li, Jianhai Huang Institute of Refrigeration and Cryogenics, Shanghai Jiao Tong University, Shanghai 200030, P.R.China
Boiling heat transfer characteristics and data on sintered porous layer surface are summarized. Non-dimensional criteria which affect the pool boiling and channel boiling are analyzed. Then non-dimensional criteria correlation of pool boiling and channel boiling which offer designing foundations for industrial application of this technique are obtained
INTRODUCTION There are two methods to increase boiling heat transfer, 1)enlarge the heat transfer area, 2)enhance the boiling heat transfer coefficient. The latter doesn't involve consumption of material, which is an active and effective way. It's effective to improve nucleate boiling heat transfer coefficient by means of porous layer surfaces. Boiling heat transfer coefficient can be raised up to approximately 10 times as large as that of a smooth surface. The critical heat flux can be raised up to 80%[2] more than that of a smooth surface. It's because of the foregoing merits that its manufacturing technology and technique develop quickly. However, boiling heat transfer is a liquid-gas two phase process, whose heat transfer and mass transfer are extremely complex, then, there is two-phase flow of capillary in the porous layers , which make the boiling heat transfer mechanism even more complicate. Much research have been carried out by over a semi-century. Mechanisms were proposed according to the range of research respectively, which fail to agree with each other. Currently main research methods can be classified into two groups:(1) set up a mechanism model on the basis of porous surface structures combined with experimental investigation. (2) find out the dimensionless criteria that affect the heat transfer based on experiment and propose dimensionless heat transfer correlation, which largely aim at respective research range. This article is intended to summarize dimensionless criteria correlation applicable to boiling heat transfer on porous layer surfaces on the basis of a series of experiments. Boiling Heat Transfer Characteristics of Sintered Porous Surface Summarizing past views and our results of experiments, we consider that the enhancement of boiling heat transfer on sintered porous surface lies in the magnanimous interconnected large-sized hollow pores within the porous surface. These pores can effectively hold back the vapor to form nucleate centers, which enables them to sustain nucleate boiling under the condition of relatively small superheat. The foregoing ways are not possessed by smooth surface. Therefore, heat transfer on porous layer surface is enhanced 5 to 10 times that of smooth surface. Effect of Main Factors on Nucleate Boiling on Sintered Porous Layer Surface Heat transfer characteristics of porous layer surface are decided mainly by structural characteristics, physical properties of materials and boiling medium. The main reason that makes heat transfer on porous layer different from that on smooth surface is from structural characteristics. Structural characteristics of porous layer are expressed by particles diameter dp, thickness of the porous layer 6 and the porosity e. There exists an optimum particle diameter with which boiling heat transfer coefficient reaches maximum when thickness of the porous layer and the porosity are constant. This is found to agree with our analysis and experiments[4]. As the particle diameter decreases, the overall surface area in porous layer increases, 581
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which leads to the enhancement of boiling heat transfer, while the liquid-gas flow resistance in the capillary increases at the same time, which is not beneficial to boiling heat transfer. Likewise, there exists an optimum thickness of the porous layer with which boiling heat transfer coefficient reaches maximum when the porosity and particle diameter are set. It is the same reason as that of the foregoing. Literature[3]'s conclusion is aimed at optimum particle diameter of a definite boiling media matched with thickness of porous layer. This is in line with the above analysis. Control of the porosity is relatively difficult, and there's few experimental investigation on it. There's an optimum porosity based on the analysis of heat transfer area and flow resistance inside capillary pores. Currently, the porosity between 40% and 60% is considered proper. C O R R E L A T I O N OF POOL BOILING HEAT TRANSFER
There are many interconnected factors that affect the nucleate boiling of sintered porous layer surface. How to properly select criteria affect the accuracy and creditability of criteria correlation directly. Though there are many public reports on dimensionless criteria correlation, few can undergo further tests. Some indexes produce essential change when data are reduced or conflict with qualitative mechanism model. Therefore, we should find a dimensionless correlation in accordance with it's physical significance and correlative index don't produce great and essential changes when the numbers of correlative data reduce and increase. Dimensionless heat transfer correlation of cryogenic fluid nucleate boiling on smooth surface are summarized as below[7].
Nu - A,(r, e)(Gr pr)m[E,g(t')] n E 1=
Hs~pg A T
there, A,(r,O) is a coefficient related to surface condition and physical properties of fluid surface, g(P) is a coefficient related to system pressure while p=lbar, ~:(P)-1.
m,n obtained through correlating
experimental data. The equation expresses briefly the criterion GrPr(movement and physical properties of liquid) and the vapor criterion El(liquid disturbance and the latent heat transfer caused by vaporization on heating surface). With reference to above equation, the criteria which affect nucleate boiling on sintered porous layer surface are as follows: 1 Re and Pr which express movement and physical properties of fluid should be exhibited separately, because of the higher frequency of vapor bubbles escaping from porous layer surface, smaller departure diameter and greater agitation than those of a smooth surface. 2 A~(r,O) is used to express heating surface characteristic of a smooth surface. While the porosity, the thickness of the porous layer and particle diameter are major characteristic parameters of porous surface, but, best correlating result is obtained by adopting ~-,6/ dj, as criterion. 3 (1- C)2p /(a',;[,) is expressed as criterion for taking the peculiarity of porous layer structure into account and introducing the effect of thermal conduction of sintered particles. 4 p~ / p~ is used to express the effect of pressure. Thus dimensionless criteria correlation applicable to nucleate boiling on sintered porous layer surface is as below
Nu - "f( Re" pr" ~'6 (l - e')2 v P' I
Ic) was developed. The transient heat transfer from the Ag shea.tl~ed ta,pe in liquid nitrogen was measured for the exponential heat inputs with the period, r, ra.nging from 10 ms to 10 s by using the newly developed temperature estimation method for the tape. Transient non-boiling heat transfer coefftcients for the tape was confirmed to be in good agreement with the corresponding theoretical values dependent on 7-. The q~, for r on the tape can be classified into two groups: that for FDNB and that due to direct transition. The direct transition from non-boiling to film boiling without nucleate boiling was observed for 7- _ 900 ms. The q~, due to direct transition is significantly lower than that. for FDNB" the q~, for the periods of 900 ms and 10 Ins, for instance, are about 15 % and 40 % respectively, of the steady-state CHF. REFERENCES
1. Sinha, D.N., Brodie, L.C., Semllra, J.S., and Young, F.M., Premature transition to stable fihn boiling initiated by power tra.nsients in liquid llitrogen, Cryogenics, (1979) 19, 225-230 2. Tsukamoto, O., and Uyemura, T., Observation of bubble formation mechanism of liquid nitrogen subjected to transient heating, A(lvances in cryogenic eng. 25, (1980), 476-482 3. Shiotsu, M., tlata, K., and Sitkurai, A., lieterogeneous spontaneous nucleation temperature on solid surface in liquid nitrogen, Advances in cryogenic eng. 35, {1990) 437-445 4. Sakurai, A., Shiotsu, M., and ]lata, K., Boiling heat transfer from a horizontal cylinder in liquid nitrogen, Heat transfer and superconducting magnetic energy storage, (1992) ASME tITD-Vol.211, 7-18 5. Sakurai, A., Shiotsu, M., and Ha.ta. I,:., Ne.w tra.nsition phenomena to fihn boiling due to increasing heat inputs on a solid surface in 1)ressurized liquids, Instability in two-phase flow systems, (1993) ASME HTD-Vol.260, 27-39 6. Sate, K., and ttikata, T., Critical currents of superconducting BiPbSrCaCuO tapes in the magnetic flux density range 0-19.75 T at 4.2, 15 and 20 I(, Appl. phys. lett., (1990) 57, 1928-1929 7. Sakura.i, A., and Shiotsu, M., Transient pool boililig heat transfer, part 1, incipient boiling superheat, ASME J. tteat Transfer, (1.(_)77) 9i), 547-553 8. Takeuchi, Y., Hata, K., Shiotsu, M., a.~t(l Sa.kurai, A., A gener;d correlation for natural convection heat transfer from horizontal cylinder i~ liqltids a.J~d gases, General l)a.l)ers in heat transfer, (1992) ASME HTD-Vol.204, 183-189
APPENDIX
C o r r e l a t i o n for N o n - b o i l i n g H e a t T r a n s f e r d u e to E x p o n e n t i a l H e a t I n p u t s Sakurai and Shiotsu [7] rel)orted that the nonl)oiling ]lea.t transfer coefficients on a vertical fiat plate due to exponential heat inputs, Qo ct/r, wil/t tile l)eriods 7- shorter than 100 nts c~tn be described by the theoretical values derived from thermal coltduct.ioz~ eq~la.lioll for the liquid, and those for r longer than 1 s by the theoretical values for natura.l colivection ltt,.at trallsfer. The conduction heat transfer coefficients, h,.:, a.r(: M)proxintately given by the following equation at a time longer than t = 3r after tlle initia.lion of tl~e heat inl)ut [7].
(kzpt%z/r) 112
hc-
(5)
Laminar natural convection heat transfer coefficients, b,,,, on a vertical plate with the height H for wide ranges of Rayleigh and Pra, ndtl lllllnbe.rs are exi)rcsscd by the following equation [8]. hn = 1 . 2 ( k l / t i ) x
10 y
(6)
where Y = 0.193.185 + 0.1,t50:37X + (I.6G.132:~IO-"X 2 - 0 . 2 3 9 , t 3 9 1 0 - 3 X 3 - 0.23861310-4X 4, R.f = G r*t:'rz/(,l + 91"r 1/~ + 10P/') .
X = loglo(R.f),
Non-boiling heat transfer coetficient.s for the il,t(:rlne(liate region, h,,~, (1 s > 7- > 100 ms) can be expressed by the following equation[7]. -
+ h,",.
(7)
Surface T r e a t m e n t of A l u m i n u m Heat S w i t c h
Toshinobu Shigematsu, Minoru Maeda, Masatoshi Takeshita, Yoshiko Fujii, Masaki Nakamura, Minoru Yamaguchi, Toyoichiro Shigi and Hiroshi Ishii* Dept. of Applied Phys., Okayama Univ. of Science, Ridaicho 1-1, Okayama 700, Japan *Okayama Ceramics Research Foundation, Nishikatakami 1406-18, Bizen, Okayama 705, Japan
Although aluminum is the most promising as a material of the superconducting heat switch at ultra low temperatures, it has not been widely employed so far. The reason is that it is very difficult to ensure the metallic contact between the aluminum strip and the copper holder. In order to realize such a condition, sputtering and plating of gold on the aluminum surface were tried after mechanical polishing, chemical etching and chemical substitution. Evaluation was done by measuring the electrical contact resistance between the aluminum specimen and the copper holder at 4.2 K. The gold-plated specimen and the gold-sputtered specimen, both treated in the Bonder-dip solution beforehand,showed the same contact resistivity, 5 n ~/cm ~-, the lowest value ever reported.
INTRODUCTION It is no exaggeration to say that at ultra low temperatures the heat switch has a decisive influence on the experiments, especially on the specific heat measurements.
In this temperature region, a metallic
superconductor with the transition temperature higher than 1 K is usually employed for the heat switch.
In
the superconducting state, the heat conduction is very poor, governed by the lattice heat conduction, because the Cooper pair does not carry entropy.
By applying the magnetic field larger than the critical
field, the heat switch restores the metallic thermal conductivity. Aluminum has the high Debye temperature and has no isotope, the former corresponds to the low lattice thermal conductivity at low temperatures and the latter to the high electronic thermal conductivity also at low temperatures. materials.
Therefore, aluminum is the most promising as a material of the heat switch over the other
However, it has not been widely employed so far from the following reason.
Aluminum is
covered with the hard oxide layer which is difficult to be taken off completely and the cleaned surface is easily re-oxidized if the surface is exposed to the air even for a short time.
This oxide layer on the
aluminum surface makes the thermal contact resistance very large. In order to overcome this problem, we tried gold-sputtering and gold-plating On the aluminum surface after removing the oxide layer by using various procedures. 621
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EXPERIMENTAL PROCEDURES The aluminum specimen of high purity (6N) had a shape, 0.1 mm thick, 10 mm wide and 20 mm long,
and its both ends were
gold-sputtered or gold-plated in 5 mm length on both surfaces. For evaluation of the fabricated specimen, it was mounted on the gold-plated copper holder as shown in Figure 1.
The electrical
contact resistances at both ends including the specimen resistance were measured at 4.2 K With four terminal method by increasing the current from - 10 A to + 10 A.
In order to
determine the specimen resistance, two
Figure ! Deviccto measure the electrical contact resistivity.
voltage leads were attached on the surface of the specimen with silver paste. This value was - 5 n Q for all specimens. I. Gold-sputtering after simple chemical etching The gold-sputtering method was adopted at first from the following reasons" (1) Special techniques are not necessary. (2) The surface of the specimen has a possibility to be cleaned by anti-sputtering before gold-sputtering. (3) The specimen is annealed in the course of sputtering. Before gold-sputtering,
the following chemical treatments were performed in the cell vibrated
supersonically and in the nitrogen atmosphere" Wash in the acetone ~ Dip in the 50% HNO, solution.
Dip in the 50% NaOH solution
After anti-sputtering, gold was sputtered on the aluminum surface at 5 kV
target voltage and 0.5 mAJcm 2 target current density in the 0.5 Pa Ar (6N5) atmosphere flowing at a rate of 10 atm 9cc/min. Although many trials were done by changing the time of chemical etching, that of anti-sputtering and that of gold.sputtering, the contact resistivity was only 10.0/z Q/cm 2 at best. We thought that chemical etching was too weak for taking off the hard oxide layer on the aluminum surface. 2. Gold-plating after mechanical polishing and chemical etching In order to remove the oxide layer on the aluminum surface more effectively, the No. 1 specimen was fabricated with the same procedure as Mueller et al.lll as shown in Table 1.
In this case, gold was
electro-plated just as they did to compare the both results. The contact resistivity of this specimen was reduced to about a half of the previous value but still very large, and the plated gold could be easily removed by the soft touch of a finger. The preceding experiments showed difficulty of taking off the strong oxide layer with chemical treatments, we decided to polish the aluminum surface at first with #3000 sandpaper coated firmly with 10/z m aluminum oxide powder. treatment.
The contact resistivity of the No.2 specimen was reduced one order with this
For the No.3 specimen, the acid solution was changed to stronger one.
on the No.3 specimen was not removed off so easily.
The plated gold layer
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Table 1 Specimen
Treatmentprocedure
2
';Contact resistivity at 4.2K
O. 6 /1. f ) / c m 2
:(~)--+(~)--+(~---+@--+(~,- - ~ ( ~ + @ ~ @ ~ @
3
0.5
u n / cm 2
(~ Polish the aluminum surface with #3000 sandpaper. (~) Wash in the acetone. (~ Wash in the alkaline cleanerat 75 0(2 for 60 see (22 ~ fl of Na3PO 4 912H20 and 22 ~ 12 of Na2CO3). @ Dip in the acid bath for 15 see (50% HNO3). @ ' Dip in the stronger acid bath for 15 see (equal volume mixture of HF, HNO 3 and water). (~) Dip in the zincate solution at (22-t-2) ~ for 60 see ( 1 g/~ of FeC13 96H20, 100 g/12 of ZnO, 525 g/12 of NaOH, l0 ~ 12of C4H4KNaO o 94HeO). (~ Dip in the same acid bath as @ for 30 sec. (~' Dip in the same acid bath as @ ' for 30 see. (~ Dip in the same zincate solution as (~ at (22-1- 2) ~
for 10 sec.
(~ Copper strike; at 26 mA/cm 2 for 2 min and then 13 mA/cm 2 for 2 min using copper anode in the following so|ution; (41.3 g/t2 of CuCN, 50.8 ~ 12of NaCN, 30 ~ 12of Na.CO~, 60 g/12 of C4H4KNaO,- 4HeO). (~) Gold plate; deposit 1/x m at 0.5 mA/cm 2 for 4 min. 3. Gold-plating after treatment in the Bondar-dip solution Today, effective zincate solutions are sold for industrial use.
So, we decided to employ the Bondar-dip
solution [2] on the market in place of the zincate solution shown in Table 1.
Accordingly, the processes
(~ and (~ were replaced by the process of dipping in the Bonder-dip solution ((~)). The results are shown in Table 2. The No.6 and No.7 specimens give the best result, 5 n ~ / c m 2, the smallest contact resistivity ever reported [3].
This fact may indicate that the oxide layer on the aluminum
surface was completely removed. Table 2 Specimen
Treatment pr(x:edure
!
5
.
.
.
.
.
|174
!
2
@--+(~)--+(~)--+(~)'--+(j~(~'7~~~
--,|
6 .
Number of peffonrtance of Contact resistivity' the substitution process at 4.2K
.
.
.
7
.
.
.
.
'
.
.
.
.
.
.
.
.
.
.
.
.
.
.
.
.
.
.
.
.
2
--, -+|174 .
.
/
t / / -,b. . . . . . . . . . . . . . . . . . . . . . .
(~--+(~)--+(~)--+@'~(~)--,(~~~~
O. O15/.1. f ) / c m 2
3 !
'
0 . 0 0 5 u o/
. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
4
0
.
m
~. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
005 # f ) / c m 2
@ Dip in the Bondar-dip solution Comparing Tables 1 and 2, it can be said that the Bondar-dip solution forms a better substitute layer than the zincate solution. substrate.
In the case of the zincate solution, the substitute layer tends to form a thick and porous
While, in the case of the Bondar-dip solution, the layer precipitated in the early stage of the
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substitution contains relatively large amount of copper.
This layer covers the aluminum surface uniformly
to make strong bonding with aluminum and serves to suppress rapid growing of the substitute layer to form a thin and dense substrate. 4. Gold-sputtering after treatment in the Bondar-dip solution At the end, gold-sputtering was performed again. The chemical procedure before gold-sputtering was the same as shown in Tab|e 2. The results are indicated in Table 3. In the case of the No.8 specimen, the substitute layer was dissolved out by the acid solution just before sputtering. But, for the No.9 specimen, gold was sputtered on the substitute layer.
From comparison of
the contact resistivity of these specimens, it can be inferred that the oxidized layer formed during the setting to the sputtering apparatus could not be removed by anti-sputterino~ . The No. 10 specimen has the smallest contact resistivity, 5 n Q / c m 2, just the same as the No.6 and No.7 specimens. Table 3 Spec~en
Treatment procedure
Anti-sputtering time Sputteringtime
Contactresistivity at 4.2K
8
5 min
5 min
0.095 # f'//cm 2
9
5 min
5 min
0.025 # f ) / c m e
5 min
5 min
0.005 # f ) / c m 2
CONCLUSION When aluminum is used for a superconducting heat switch, it is important to reduce the thermal contact resistance between the aluminum strip and the copper holder by removing the oxide layer on the aluminum surface.
We have developed the best procedure to fabricate an aluminum heat switch by applying
gold-sputtering or gold-plating on the aluminum surface after mechanical polishing, chemical etching and chemical substitution with the Bondar-dip solution. Both gold-sputtering and gold-plating gave the similar result, 5 nQ/cm e in electricalcontactresistivity, which is the lowest value ever reported. REFERENCES 1
Mueller,R.M.,Buchal,C.,Oversluizen,T. and Pobell,F., Superconducting aluminum heat switch and plated presscontacts for use at ultralow temperatures, Rev.Sci.lnstrum. (1978) 49 No.4 515-518
2
Bondar-dip solution, CANNING JAPAN Co. Ltd., Kameido I-8-4, KBt~-ku, Tokyo 136, Japan
3
Bunkov,Yu.M., Superconducting aluminum heat switch prepared by diffusion welding, Cryogenics (1989) 29 September938-939
The Study on the Solid Thermal Contact Resistance at Low Temperatures
Xu Lie*,
Zhou Shuliang*,
Yang Jun*,
Xu Jiamei**
*Low Temperature Center, Shanghai Jiao Tong University, Shanghai 200030, China **Shanghai Sunny Research Institute of Environment and Energy, Shanghai 200040, China
This paper mainly studies thermal contact resistance (TCR) of solid interfaces. On the basis of Reedwood-Williamson model, a thermal contact resistance model is proposed. Experiments and analyses of TCR of stainless steel and aluminum have been done which include the effects of pressure, temperature, roughness, stuffing and different materials.
INTRODUCTION In many advanced research and application areas of high technology, it is necessary to keep the measuring equipment or a system in a low temperature environment. Heat has to be taken from it through a passage connected to a cooler or low temperature liquid. The contact resistance of the solid surface will affect the heat transfer performance of the system. Lower values of TCR can improve the performance of heat transfer, while higher values can be used for insulation[ 1]. In the experiments, effects on TCR of pressure, temperature, surface roughness and material properties have been studied, some of the results are presented in this paper which may be useful for the engineering design.
EFFECT OF SURFACE ROUGHNESS When heat is transferred from a solid surface to another through an interface, usually there exits an excess temperature drop due to the interface A T - T~ - ~
(1)
Defining R - q / A T as contact resistance, which is composed of thermal resistance Rs of solids, thermal convection resistance Rf of fluids between the contacting surfaces and thermal radiation resistance R,. of the two surfaces:
R
-
R.,. + Rj~ + R,.
(2)
If the heat transfer occurs in vacuum, the fluid heat transfer can be treated as conduction heat transfer through fluid, the thermal contact resistance is R f - Kt-Aj./6, where 6 is the average thickness of the gap between two surfaces. Compared to solid surface heat conduction and fluid heat transfer, the radiation effect is very small and can be neglected. Actually, the percentage of solid surface thermal resistance or fluid thermal resistance to the total resistance is changing due to the effects of thermal conductivity of the solid and fluid, and also the actual contact areas. The fluid (mainly gas) is under vacuum, its heat transfer effect can be calculated by the formula of molecular flow[2]. So in this paper the effect of solid surface heat conduction will mainly be analyzed. The actual contact area is only part of the total area, so the heat flux will have a constriction at the joint. Under the conditions of constant material properties and no heat source, just like fluid flowing through a hole must obey the equation of the conservation of mass, the equation of conservation of energy is: 625
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ICEC16/ICMC Proceedings
+
+
-0
(3)
here z is the axial distance from contact. There are many small channels on the contact surface, the above equation for one channel model can be written using cylindrical coordinates as: 02 T
1 (7/'
~'~
r ~"
C 2T
+
-0
-~-
(4)
according to the boundary condition, when Z --~ oe
; r - c,, + 9
/
z
(s)
when there is no insulating gap between the interface, 71:=. - C., actually, the average temperature at the interfaceis
:
~ ,-, - C,., + q -=
x ~(a')
4ak
(6)
from Equation (5) and (6), the temperature drop due to the constriction of heat flux on the contact is: A T - TI_=,:' -7~:=,:,--
4 aqk
x ~(s)
(7)
~J is the function of shape and size of the contact area. From the analyses above, it can be seen that thermal contact resistance of solid interface is mainly caused by the constriction of heat flux . Practically, there are many small areas connected together on a interface, their sizes and distribution are very complicated. For this kind of surfaces, we give the formulae below on the basis of analyses[3 ].
R-1
or,.' 3D. K
•
E
F
•
V(,5.)
~,_(d/cr. '
. .
h - C(F). D ~ - -
o- ["'"
(8)
(9)
Formulae above indicate that TCR is in inverse relation with the heat conductivity of materials and the density of rough granule on the surface, proportional with surface roughness and the modulus of elasticity of the materials.
RESULTS AND ANALYSES OF THE EXPERIMENTS Experimental results of TCR of aluminum and stainless steel with various conditions are shown in Figure 1 to Figure 5. Effect of Loads Figure 1 shows the thermal conductance data of aluminum pairs (No.1 to No.5) at 135k with different surface conditions under loads from 1MPa to 3MPa. They have the same tendency vs. pressure though the values are not same.
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Effect of Temperature Figure 2 is the thermal conductance of stainless steel pair No. 1. This figure shows that thermal conductance increases with increase in interface temperature. Published data of thermal conductivity of bulk material has been used to calculate the heat flux, therefore, the effect of temperature on interface thermal conductance is influenced partly by bulk material thermal conductivity. This data has a big difference with reference [4]. In reference [4], when the average temperature of the contact was changed from 30~:C to 190~C, thermal conductance has changed four times. But we got only 25% change when temperature changed from 130K to 350K. The discrepancy can be due to the differences in the experimental conditions. The data in reference [4] was obtained at the ambient pressure with foam type insulation materials, so thermal conductance of the gas between the interface and also the insulating material would affect the results, while in the present work data was taken under the conditions of high vacuum and with the protection of the radiation shield at lower temperatures.
Figure 1 Thermal conductance versus pressure of A1-A1 interface with different surface conditions (No. 1-No.5) at 135K
Figure 2 Thermal conductance of stainless steel interface at different temperatures (P 1- 1.06MPa, P2 = 1.3 6MPa, P3 = 1.67 MPa, P4=2.09MPa, P 5=2.86MPa)
Figure 3 Thermal contact resistance of stainless steel with different surface roughness at 350K(P 1= 1.06MPa, P2 = 1.3 6MPa,P3 = 1. 67MPa, P4=2.09MPa,P 5=2.86MPa)
Figure 4 Thermal contact resistance at different temperatures under ambient pressure
Effect of Surface Roughness Figure 3 shows thermal contact resistance of stainless steel samples versus surface roughness at 350K. At certain temperature and pressure, thermal contact resistance will increase with the increasing of the
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roughness. At 2.86MPa, when the roughness was reduced from 44/z m to 18 ~z m, thermal contact resistant lowed by 40%, while the roughness changed from 18 g m to 2 tz m, thermal contact resistance decreased by 80%. Effect of Stuffing in Gap The value of thermal conductance of stainless steel interface under 1.06MPa augmented with vacuum grease has been measured as 2473W/mZ'K, which is 7 times better than the data without the augmentation of vacuum grease. The results shows that augmentation of vacuum grease can reduce the TCR effectively. Comparison Between Theoretical Analyses and Experiments The experimental data and the references imply that thermal conductance has a linear correlation with pressure. To a certain sample, at a certain temperature Equation (9) can be rewritten as h=c-F. So theoretically, when the pressure of contact is 0, thermal conductance will also be 0, but we find that there exists a residual thermal conductance. As shown in Figure 5 , thermal conductance h=h0+c.F, this is caused by some other heat transfer mechanism not included in the analyses, and the presumption of fully elastic deformation. So it is better to include the residual thermal conductance to calculate thermal conductance. Fitted formula for thermal conductance of stainless steel shown in Figure 5 is: h=221+267P (350K) Figure 5 Relation between TCR and pressure of stainless steel at 3 50K CONCLUSIONS Theoretical analyses agree with the experimental results qualitatively. Pressure, temperature, surface roughness, material properties and filling the gap will affect thermal conductance as follows 1. the pressure applied on the contact surfaces affect the thermal conductance, and it can be expressed as h=h0+c.F. 2. thermal conductance increases with the rising of the temperature. 3. surface roughness has a large effect on thermal conductance, thermal conductance will increase with reducing of surface roughness, especially at the low roughness condition. the properties of the bulk material will affect thermal conductance of the interface. Filling heat conducting medium into the gap can increase thermal conductance. So it is possible to increase or decrease thermal conductance by changing pressure, roughness, filling the heat conducting grease, and plating with soft heat conduct film, etc. ~
REFERENCES 1. 2. 3. 4.
Snaith, B., Probert, S.D., Callaghan, P.W., Thermal resistance of pressed contacts, Applied Energy (1988)20 31-84. Xu, L., Fang, R.S., Ma, Q.F., Insulating Technology, National Defense Industry Press, China, (1990) 210-219. Yang, J., Research on the Heat Contact Resistance at Low Temperature, thesis for Master's degree, Shanghai Jiao Tong University, China (1995). Gu, W.L., Experimental research of heat contact resistance, Journal of Nanjing Aviation Institute (1992)24 44-53.
Experimental Study on Thermal Contact Conductance at Liquid Helium Temperature Kazumi Sunada,
Yoon-Myung Kang
MEC Laboratory, DAIKIN INDUSTRIES, LTD., 3 Miyuldgaoka, Tsukuba 305, Japan When cooling an object in a cryogen-free, cryocooled system at liquid helium temperature, it is important to quantitatively estimate the thermal resistance that occurs at the contact surface between the cooling stage and the object. Here we employed two methods to estimate this thermal contact conductance at liquid helium temperature : (1) an empirical equation and (2) the Wiedemann-Franz law. We compared the estimated conductance with that measured at liquid helium temperature, both methods were proved valid. Additionally, we measured the thermal conductivity and the electric resistance of phosphor bronze screen stacks at room temperature and at low temperature.
INTRODUCTION When cooling an object in a cryogen-free, cryoeooled system at liquid helium temperature, it is important to quantitatively estimate the thermal resistance that occurs at the contact surface between the cooling stage and the object. A recent comparison shows a large discrepancy in the published values for the thermal contact conductance of nominally similar contacts at room temperature; sometimes differing by more than six orders of magnitude. Furthermore, little data is available on thermal contact conductance measured at liquid helium temperature of bare contacts. Therefore, there is a neeA for bare contacts. Here, we measured the thermal contact conductance of bare contacts at liquid helium temperature. First, as reference, we measured the thermal contact conductance of bare contacts made of copper (OFHC) at liquid helium temperature. We then compared these measurements with estimates calculated using an empirical equation derived by P.W. O'Callaghan and his colleagues [1] for the thermal contact conductance at room temperature under a vacuum in which they assumed elastic contacts. Second, we estimated the thermal conductance of contacts with and without interposers at liquid helium temperature by measuring the electrical resistance of the contacts at room temperature. For electrically conductive contacts consisting of conductive blocks and interposers such as indium foil, we estimated the thermal conductance at liquid helium temperature (CT4K) from the Wiedemann-Franz law, taking into account the temperature dependency of the contact's properties. For reference, we measured the thermal conductance of contacts with a non-conductive thermal grease (Apiezon-N). We then compared these estimates with our measurements, showing that the thermal contact conductance of electrically conductive contacts made of copper at liquid helium temperature could be estimated from the electrical resistance at room temperature within an accuracy range of about 30%. Additionally, we measured the electrical resistance of phosphor bronze screen stacks, which are used in regenerators of cryocoolers, at room temperature and low temperature. We used the measured thermal conductance of the stacks at low temperature to calculate the thermal conductivity of the stacks. Because of the result of this, the Wiedemann-Franz law did not hold between the electrical resistance of the stacks and the thermal conductivity of them. EXPERIMENTAL APPARATUS We chose copper (OFHC) as our sample because k is often used in cryoeoolers, and because the thermal contact conductance of copper has been measured by many researchers at room temperature. The contact specimen was a copper (OFHC) cylinder with a nominal cross-section of 1.4 cm 2, an outer diameter of 14 mm, and an inner diameter of 5.0 mm. To normalize the data as the following equation (1), we measured the surface roughness before and after each experiment. A load was applied to the contact specimen by a stainless steel bolt at pressured from 0 Mpa to 20 MPa and was measured by strain gages. A steady-state longitudinal heat flux method was used to measure the thermal contact conductance. To occur heat flow, a button heater was placed on the upper samples. To monitor the specimen temperature, Ge thermometers were installed in both the upper and lower samples such that the thermometer axes were parallel to and located 5 mm from the contact interface. 629
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The electrical resistance of the contact was measured using a 4-wire DC method. The wires were soldered to both the upper and lower samples. To keep the sample heating negligible, current was controlled by the thermometer response. We thus measured the thermal contact conductance of copper (OFHC) without a thermal interposer (i.e., bare) with indium foil, and with Apiezon-N grease. For the indium foil, we placed a thin donut of indium foil between the contact interfaces, in which the outer diameter of the donut was 14 mm, the inner diameter was 5.0 mm, and the thickness 0.1 mm. For the Apiezon-N, we applied the grease to both contact surfaces, provided the contact was electrically conductive. EMPIRICAL EQUATION FOR THERMAL CONDUCTANCE AT ROOM TEMPERATURE We normalized the load and the thermal contact conductance with surface roughness etc. using the results from following equation by P. W. O'CaUaghan et al. [ 1] as reference.
W*= W
a2H
C*=-.A~ aX
(1)
where W* is the normalized load applying the contact, C* is the normalized thermal contact conductance, W is the load, C is the thermal contact conductance, 6 is the surface roughness, H is the effective elastic modules, An is the nominal cross-sectional area, 2Lis the thermal conductivity, of the sample. By assuming the contacts were elastic contacts, O'CaUaghan et al. derived an empirical equation (2) from previous experimental data (344 points) that included data for contact between various metals ( e.g., copper-copper, aluminum-aluminum, etc.) in a vacuum at room temperature
Figure 1
Comparisons of normalized thermal contact conductance
Hgure 2 Compafis.ons.gf t.he.rma.1conductances measured at llqU1(1 llellum temperature with and without Indium foil or APIEZON-N and those calculated from electric resistancesat room temperature
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631
(2)
Figure 1 shows the measured and empirical thermal contact conductance for copper (OFHC) contacts without interposers at low temperature. [2,3] The agreement between the two confh'rns that the Callaghan empirical equation can accurately explain this thermal contact conductance. EFFECT OF INTERPOSERS AND ELECTRICAL RESISTANCE OF CONTACT The Lorentz Number of most metals depends on the temperature. [4, 5] For copper (OFHC), the Wiedemann-Franz law holds near liquid helium temperature and room temperature. Using this law, Lee. et al. [4] explained the relationship between the thermal contact conductance (CT) and the electrical contact resistance (Rel) as CT=
LT AnRel
(3)
where L is the Lorentz Number and T is temperature. Using copper (OFHC) as the contact specimen, we measured the thermal contact conductance at liquid helium temperature. And the electrical contact resistance was measured at room temperature and liquid helium temperature. The relationship between the thermal contact conductance at liquid helium temperature: CT4K and the electrical contact resistance at liquid helium temperature : ReMK can be explained by the equation (3). Therefore, if we can derive Rd4K, we can calculate CT4K. When RelnK was too small to measure, we extrapolated the value by assuming the pressure that was decreased by cooling. Figure 2 shows the comparison of CTnK calculated using equation (3) for the electrical contact resistance at room temperature with that measured at liquid helium temperature. These two values agree well. MEASUREMENT OF PHOSPHOR BRONZE SCREEN STACKS Additionally, we measured the thermal conductivity of phosphor bronze screen stacks at low temperature and the electrical resistance of them from 4K to 300K. These stacks are used in the regenerators of cryocoolers. Each screen was 25.4 mm in diameter and was a 200-mesh composed of 53 Ixm diameter wire. We measured stacks of 40 and 100 screens. After measuring the thermal conductance of a stack at low temperature, we calculated its thermal conductivity in bulk shown in Figure 3. And condition 1 is a stack that was cleaned ultrasonically before measurement. Condition 2 is a stack that was cleaned ultrasonically before measurement and then degassed 8 hours at 393K. Condition 3 is a stack that was cleaned ultrasonically before measurement and then oxidized 8 hours at 393K. Figure 4 shows the Lorentz Number of the stacks calculated from the thermal conductivity of the stacks and the electrical resistance of the stacks using Wiedemann-Franz law in equation (3). -At,xlO0
W-2-d
[%]
(4)
where d is wire diameter of phosphor bronze consisting screen mesh, At is thickness of a clamped screen, and W is the degree of clamping of the screens. The pressure applied the contact at low temperature was measured from 0.01 to 0.4 MPa by the strain gage. The Lorentz Number of phosphor bronze bulk was about 2.9 * 10* W f g K 2 at 300K. In contrast, that of the phosphor bronze screen stack was large more than 2 orders of magnitude compared to 2.9 * 10.8 WDJK 2, provided the Lorentz Number of the stack decreased as the pressure was increased. For tI' was 104 %, the Lorentz Number of the stack for condition 1 was around 70 * 10"8 WD/K 2 at liquid helium temperature. Further work is needed to determine the causes for the increase in the Lorentz Number of the stacks, namely, measurements in which (a) higher contact pressure is applied and (b) the contact surface is cleaned chemically. The result for condition 3 indicates that an oxidized contact surface strongly influences the Lorentz Number; this influence agrees with that reported in a previous study by Nilles, et al. [2] Therefore, oxidation of the contact surface was a factor in increasing the Lorentz Number of the stacks.
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Figure 3 Comparisons of thermal conductivity for phosphor bronze stacks
Figure 4 Lorentz Number of phosphor bronze stacks
CONCLUSION The equation (2) also helps us to quantitatively understand the thermal conductance of bare contacts at liquid helium temperature. For electrically conductive contacts consisting of electrically conductive blocks and interposers such as indium foil, the thermal conductance at liquid helium temperature could be estimated with WiedemannFranz law when the temperature dependency of the properties are considered. In contrast, the thermal conductance of phosphor bronze stacks whose surface was oxidized could not be estirnated with Wiedemann-Franz law between 4K and 300K. REFERENCES 1 2 3 4 5 6
O'CaUaghan, P.W. and Probert, S. D. 9Journal Mechanical Engineering Scienc~ (1974) 16 No.7. 4155 Nilles, M. J. and Van Sciver, S.W. 9.Adv. Cryogenic Engineering (1988) 34 443-450 Yu. J., Yee, A.L. and Schwall, R.E. 9Cryogenics (1992) 32 610-615 Lee, A.C., Ravikumar K.V. and Frederking T.H.K. 9Cryogenics (1994) 34 451-456 Clark, A.F., Childs, G. E. and Wallace, G.H. 9Cryogenics (1970) 4 295-305 Li, R. Hashimoto, T. Ohta, K. Okamoto, H. 9Proceedings International Cryogenic Engineering Conference (1988) 12 414-417
Cryogenic engineering
Gas properties
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On the Joule-Thomson Integral Inversion Curves of Helium-3, Helium-4 and Hydrogen
B-Z. Maytal* and A. Shavit** *Rafael Institute, Cryogenic Section, P.O.Box 2250(39), Haifa 31021, Israel **Department of Mechanical Engineering, Technion-Israel Institute of Technology, Haifa 32000, Israel The Integral Inversion Curve (I.I.C.) is the contour of all thermodynamic states which exhibit a zero integral cooling under Joule-Thomson (isenthalpic) expansion from a pressure down to ambient atmospheric pressure. This is an alternative and complementary presentation of the Joule-Thomson inversion phenomena. The traditional inversion curve is a differential one. The I.I.C. has a peaking temperature dependence in the plane of pressure and temperature. This is the highest pressure for which integral cooling is still possible. However, most of the gases solidify before the peaking pressure is reached. The quantum gases, helium and hydrogen comprise the exceptional group which enables the study and verification of the predicted peaking pattern of the I.I.C. The helium-4 and hydrogen I.I.C.'s are obtained through an available numerical code. The I.I.C. of helium-3 is evaluated through an advanced, sixteen parameters equation of state. By analogy to the traditional maximum inversion curve, the maximum integral inversion reduced pressures are determined to be about 40 for helium-4 and 30 for both hydrogen and helium-3.
NOMENCLATURE Cp
Cpo
h hK; hR Ah r M
n~ P
Isobaric specific heat, J/(mole K) Isobaric specific heat at zero pressure, J/(mole K) Specific enthalpy, J/mole
•Pc
R
T
Ideal gas enthalpy, J/mole
BOIL
Residual enthalpy, h - h tG , J/mole Integral isothermal Joule-
V
Thomson effect, J/mole Molar mass, g/mole Coefficients of the He 3 equation of state, i- 1... 16, Pressure, Pa
Z Greek 9 | FI
Tc Vc
635
Critical pressure, Pa Universal constant of gases, 8.314 J/(mole K) Isenthalpic temperature drop, K Absolute temperature, K Normal boiling point, K Critical temperature, K Specific volume, m3/mole Critical specific volume, m3/mole Compressibility, P . v / ( R . T ) notation Density, mole/dm 3 Reduced temperature, T/Tc, Reduced pressure, PIPe,
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INTRODUCTION The inversion of the Joule-Thomson effect was extensively and continually studied in its differential form for about ten decades. The essence of this remarkable tradition manifests itself in the differential inversion curve. It is the locus of all thermodynamic states, (P, T),
for which the Joule-Thomson coefficient
vanishes, namely, ~t = (aT/OP)h = 0. These states serve as a border of transition from the differential cooling to the heating zone. An alternative approach of treating the inversion phenomena is the integral one. The posed question wheather integral heating (ATh > 0)or cooling (ATh < 0) is examined under finite pressure drops, rather then infenitisimal from P down to zero (or ambient). The integral inversion states are
- r ( P = 0, h ) - v(P, h ) - 0
(1)
In historical perspective, the very first scientists who explored the inversion phenomena adapted the integral approach. It was Porter [ 1] (1906) that recommended and established the differential tradition replacing those days current integral approach, not by claiming of any advantage but plainly, for simplicity of treatise of the inversion phenomena. Following Porter, scientists organized and arranged their measurements and analysis in terms of the differential version. Maytal and Van Sciver [2] proposed an empirical correlation for the I.I.C. of low acentricity gases. Maytal and Shavit [3] mapped the intensity of the Integral effect forthese gases and obtained a qualitative picture, including a closed form solution for the I.I.C., applying the Van der Waals equation of state. Maytal and Pfotenhauer [4] derived the I.I.C. via the Peng-Robinson equation of state Koeppe [5] was the first that explicitly introduced the concept of the I.I.C. accompanied by a qualitative plot. He proved that the I.I.C. has a peaking shape and the state of maximum integral inversion curve, if reached, satisfies the condition, Cp = Cpo. However, excluding helium and hydrogen, all gases solidify before their I.I.C. peaking condition takes place. The intention of the present treatise would be to (a) present the I.I.C. of helium-4 and hydrogen, (b) verify the peaking condition, (c) determine the I.I.C. for helium-3, through an advanced equation of state, and (d) display the relation between the traditional differential inversion curves and the I.I.C.s.
THE EQUATION OF THE INTEGRAL INVERSION CURVE The definition of the integral inversion curve through equation 2 is equivalent to the condition
AhT-h(P=O,T)-h(P,T)=O
(2)
The residual enthalpy, h R = h - h z~;, is the deviation of the real from the ideal gas enthalpy. Since h'~J(P = O, T) = h'~;(P, T), and h R ~ 0 because of the low pressure, we get Ahy = - h l~. Hence the alternative condition to be fulfilled by the I.I.C. would be, h R = 0. Expliciting [6] h R and setting to zero, finaly leads to the the equation of the I.I.C.,
o1_
o'9
~,-
T
(3)
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DETERMINATION OF HELIUM-3 INTEGRAL INVERSION CURVE Helium-3 is the most outstanding gas within the unique group of quantum gases. It is the strongest violator of the Principle of the Corresponding States. As such, any display of the I.I.C.s would not be complete without helium-3. Driven by the same motivation [7,8,9] the differential inversion curve of helium-3 was evaluated. The hereby applied advanced equation of state [ 11 ], where X = T + 5.6906, P
/
= Z = 1+ n~ + ~ +
x
+
9 +
7
T . n 5 "4-El6 "["
" ~ + ~
v
, 3.9 4 r
"[-
T
(4)
Integrating equation 3 following the substitution of above Z and 0, the I.I.C equation, in the plane (T, p)is obtained. One may solve it in conjunction with equation 4, thus getting the I.I.C. in the commonly used (T, P) plane..
QUANTUM GASES INTEGRAL INVERSION CURVES Information about integral inversion points of helium-4 and hydrogen is directly obtained from GASPAK numerical code [12] of their specific thermophysical properties. For each temperatue, a high pressure state is requested so that enthalpy equalizes the value at zero pressure. Collection of these states comprise the I.I.C . . . . . . . . 1 displays the I.I.C.s for helium-4 and hydrogen, in addition to helium-3 as previously obtained. The differential inversion curve of helium-3 was derived from the same equation of state [9] and supported by direct measurements [10,11] of inversion points. The differential inversion curves are extracted while applying the numerical code, GASPAK of Cryodata Inc. and searching for states of = 0. These, too, are shown in Figure 1. Inversion curves of nitrogen, as a representative of low acenticity gases, are displayed for refemce and amphasis. For each gas the space between the differential and integral inversion curves are marked for better perception of the general picture.
DISCCUSION The pattern of the I.I.C.s demonstates once again the outstanding position of the quantum gases and disobidience of the Principle of Corresponding States. The most remarkable feature is the pressure peaking form of the quantum gases I.I.C. For instance, He 4, while expanding from a higher pressure than 9.51 Mpa (FI = 41.8) down to zero, will allways warm up and never cool down. Lets consider the ratio between the peaking integral and differential inversion pressures for He 3, He 4 and H2 which respectively are, 1.26, 2.45 and 3.20. One may clearly observe a general trend: the I.I.C. approaches closer to the differential inversion curve as the quantum effect of the gas becomes stronger. The two inversion curves of He 3 are indeed relatively close. The helium isotops inversion curves (both differential and integral) are firmly distinguished altough they represent the same chemical element. The gap between helium-4 and hydrogen is similar to the gap between helium-4 amd helium-3.
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REFERENCES ~
,
o
~
~
6. 7. 8. .
10. 11. 12.
Porter, A.W., "On the Inversion Points for a Fluid Passing Through a Porous Plug and Their Use in the Testing Proposed Equations of State", Phil. Mag., Vol. 6, pp. 554, (1906). Maytal, B-Z. and Van Sciver, S.W., "Charecterization of Coolants for Joule-Thomson cryocoolers", Proc. 6th Int. Cryocoolers Conf., Vol. 1, pp. 245, Plymouth, MA, Jan., 1991. Maytal, B-Z. and Shavit, A., "On the Integral Joule-Thomson Effect", Cry_ogenics, Vol. 34, No.I, pp.19, (1994). Maytal, B-Z. and Pfotenhauer, J.M., "The Integral JT Inversion Curve by the Peng-Robinson Equation of State", Proc. 10th Intersociety Cry.ogenic Sym., Houston, TA, (March 1995). Koeppe, W., "Bemerkungen zur Inversionskurve", Ka!tetechnik, Vol. 14, pp. 399-403, (1962). Walas, S.M., Phase Equilibrium in Chemical Engineering, Butterworth Press, Boston, (1985). Maytal, B-Z., "Helium-3 Joule-Thomson Inversion Curve", Cryogenics, (in press). Vortmeyer, D., "The Joule-Thomson Coefficient of Non Polar Gas Mixtures at P ~ 0. A Theoretical Interpretation of Experiments", Kaltetechnik, Vol. 18, No. 10, pp. 383, (1966). Gibbons, R.M. and McKinley, C., "Preliminary Thermodynamic Properties of Helium-3 between 1 K and 100 K", Adv. Cryo. Eng., Vol. 13, pp. 375, Plenum Press, New York, (1968). Duant, J.D., "Preliminary Thermodynamic Data for the Inversion Curve of Helium-3", Cry_ogenic$, Vol. 10, pp. 473-475, (Dec., 1970). Kraus, J. et al., "Enthalpy-Pressure (H-P) Diagram of He3 in the Range 1.0 K> 7/Vt .
For a typical Si-diode in liquid Nitrogen, with a diode voltage of Vd = Is "R1 factor close 1.6, this leads to an n-value between 90 and 100.
(2)
=
1 V and with an ideality
The passive circuit as described here is particularly useful as a V-I simulator for testing equipment that measures the V-I characteristic of superconductors. Compared to the existing V-I simulator, operating at room temperature the voltage transition is much steeper. The V-I simulator operated at room temperature has an n-value of 27 [1]. A1 the major problems regarding the temperature changes in the room temperature device are minimised when V-I simulator is operated in a nitrogen bath at 77 K. The power generated in the diodes can be restricted to a few milli-watts maximum. Such a small power is easily absorbed in the N2 bath, without causing any measurable changes in the diode temperature nor its V(/) characteristic. The power generated in R2 is also negligible, but in RI there is a significant power of Va'L. By selecting a well cooled and stable resistor for Rt this problem can be overcome too.
EXPERIMENTAL RESULTS A V-I simulator is characterised at 77 K in order to demonstrate the feasibility of the device. Two medium sized silicon rectifier diodes (1N5401-DC, rated at room temperature for 3 A) are combined with a 50 W power resistor (Rl=100 mf~). The simulator voltage is measured over a 15 mm piece of resistive wire (R2=60 m ~ ) . The combination of the resistance value for Rl and the diode voltage leads to a voltage transition at 10 amperes. The slope of the transition can be described with equation 1 which leads to an n-value ranging from 100 to 85, in the voltage regime from 0.01 to 10 ktV. The description formulated in equation 1 becomes invalid for high current values where a significant part of the current starts to flow through the diode. Above a simulator voltage of 30 laV the n-value decreases significantly due to the internal resistance of the diode Rd (not drawn in figure 1). A V(/) curve as measured on this simulator at 77 K, over more than 6 voltage decades, is presented in figure. 2. The data points are obtained at a constant current in a step-wise manner. With a careful offset correction at zero current and sensitive (nano-)voltmeter the voltage noise can be reduced to +3 nV. The measured data can be described with an ideality factor of 7/= 1.65 and a conduction resistor of Rd = 3.5 ~ for this particular set of diodes. Is
RI
IOl
ld
Re
Vs Fig. 1 The V-I simulator circuit with a double set of diodes for bipolar operation.
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Fig. 2 The measured voltage across the V-I simulator at 77 K. The solid line represents the data fit with 7/= 1.65 and Rd = 3.5 ~ .
NOISE AND SENSITIVITY. At low voltages (V < 1 ktV) an accurate detection of the voltage becomes a delicate matter. Due to the stable (and low) temperature in the nitrogen bath and the small resistor values present in the circuit, the voltage noise in this simulator is relatively low. The combination of a ground contact in the simulator circuit and a sensitive voltmeter with a direct twisted Cu-pair to the simulator sample leads to voltage noise of (+3 nV). With regard to this noise level, a constant voltage reading is observed over the entire current range from zero to the onset of the V-I transition at 9.5 amperes. A detailed view of this transition is presented in figure 3.
STABILITY AND REPRODUCIBILITY. The stability of the V-I simulator is determined mainly by the stability of the resistor R~ that carries the largest part of the current and the diode characteristic. Due to the stable temperature and effective cooling of the liquid nitrogen bath no variations in the simulator V(/) characteristic could be observed with the available accuracy in the current measurement 0.05%. This excellent stability remains when the device is thermally cycled between 77 K and room temperature. No changes could be observed after multiple thermal cycles (investigated up to 10) within the sensitivity limits of the voltage and the accuracy of the current measurement.
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Fig. 3 The measured voltage across the V-I simulator at 77 K, magnified at the onset of the V-I transition. The solid line represents the data fit.
CONCLUSIONS The performance of a passive V-I simulator for superconductors is investigated. A stable and reproducible V(/) characteristic is shown in a nitrogen bath cryostat at 77 K. The V(/) characteristic is described well with the present descriptions for Si-diodes. If the conduction resistance of the diode is taken into account, then the V(/) curve is defined accurately over 6 orders of voltage magnitude. A stable and reproducible operation is shown within the experimental accuracy (+3 nV in voltage and +0.05% in current) The slope of the V(/) curve at the transition is for low voltages determined by the ideality factor 7/of the diode and the voltage across the diode. For a Si-diode at 77 K this leads to an n-value that is close to 100. For large voltages a higher current through the slope of the V(/) transition decreases towards a constant value that is determined by the power resistor R~ and the ratio between R2 and the conduction resistance of the diode (Rd). Besides the improvement of the stability of the cryogenic simulator compared with a device operated at room temperature, there is the additional advantage that the cryogenic version has a small size and can replace a HTS-sample in the nitrogen cryostat. This enlarges the number of possible error sources that may influence the determination of the V(/) characteristic. Additional noise sources as thermo-couple noise in the current leads and wiring, common mode errors, pickup-noise, ground loops and leakage currents are included in an equipment test with a cryogenic voltage simulator. The demonstrated properties of the novel cryogenic version of the V-I simulator makes the device suitable for use as a reference sample for critical current measuring systems. Especially in round robin tests, as for example in the on-going VAMAS programme, the cryogenic V-I simulator can be used to compare the differences between the test set-ups in the different laboratories accurately. In an industrial environment a cryogenic V-I simulator can be applied as a reference sample in a quality assurance routine of the characterisation equipment for superconductors.
REFERENCES D. Aized, et al. Comparing the accuracy of critical current measurments using the voltage-current simulator, IEEE trans. On Magn. Vol. 30, No. 4, p. 2014, 1994. Tyagi M.S., Introduction to semiconductor materials and devices, John Wiley & Sons, New York 1992.
The Estimation of Critical Current Density Using SRPM and AC Methods Shuichi Koto *, Hiroshi Nakane *, Edmund Soji Otabe ** , Teruo Matsushita ** Shigeo Nagaya ** ,and Shuji Yoshizawa **** *Department of Electrical Engineering, Kogakuin University, 1-24-2 Nishi-Shinjuku, Shinjuku-ku, Tokyo 163-91, Japan **Department of Computer Science and Electronics, Kyushu Institute of Technology, 680-4 Kawazu, Iizuka-shi, Fukuoka 820, Japan ***Electric Power Research and Development Center, Chubu Electric Power Co., Inc., 20-1 Kitasekiyama, Ohdaka-cho, Midori-ku, Nagoya 459, Japan ****Central Research Laboratory, Dowa Mining Co., Ltd., 277-1 Tobuki-cho, Hachioji-shi, Tokyo 192, Japan The critical current density ( J ~ ) of Y-Ba-Cu-O sample made by the zone melting process was estimated in SRPM and AC methods. When it was measured under AC and DC magnetic flux density, J ~ obtained in the both of the methods agrees well. There is a possibility to be able to apply the SRPM method to the measurement of J c.
INTRODUCTION We have been investigating a method which can simultaneously estimate both the resistivity ( p ) and the magnetic penetration depth ( 2. ) by vectrially measuring the difference in the impedance between two circular solenoid coils; one with and the other without a rod-shaped sample conductor. (SRPM method) [1]. It is easy to measure the frequency dependence of p and 2.. On evaluating the property of superconductors, the critical current density ( J ~ ) is taken as one of the most important properties except the temperature properties of p and 2 . In order to estimate J ~, AC method (Campbell method) is commonly used [2]. In this method, ). is estimated when the AC magnetic field ( b ) is supplied for the specimen, and J ~ is calculated from the inclination between b and ). under DC magnetic field. For the SRPM method, J ~ was obtained by using the same analysis. The values of J obtained from the both methods were compared. In this paper, the effectiveness of the SRPM method for the estimation of J ~ is discussed. METHOD FOR ESTIMATING J c The process for obtaining J c by the SRPM method is shown as the flow chart in Fig.1. In the SRPM method, the impedance of the coil was vectrially measured twice at the same temperature: once when a rod-shaped sample was coaxially inserted into the solenoid coil and then it was pulled out of the coil. The differences between the real part ( A R ) and the imaginary part ( A X ) in the impedance of the coil at the different conditions mentioned above were obtained from the measurement, while the impedance change of the solenoid is theoretically expressed in [1]. A R and A X calculated by the equation are shown in Fig.2 as the map using the parameters of 2 and p .The 30-layer solenoid coil of 832 turns with the average radius of 5.31mm, length of 3mm, the specimen of radius of 2.27mm, and the frequency of lkHz were used in Fig.1. From the point on the map the parameter of 2 is obtained. In this case, it is apparent that /! consists in the range of 10-~'~ 10 -" [m]. ). at each b is detected by magnifying the range. In the AC method, AC magnetic field was given parallel to the 701
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ICEC 16/ICMC Proceedings < SRPM Method > Impress b and B
Calculation of A R , A X
Measurement of A R , A X
Map of A R , A X
Estimation
of
Vp, V e
,
I
Calculation to Obtain Property of b - 2 '
Property of b - 2 ( p )
t Jc
Jc Fig.1
Measurement
,
o~
Comparison
< AC Method >
Flow Chart
specimen under DC magnetic field. Using the voltage of pick up coil ( V~ ) and cancel coil ( V c ), the difference between V p and V ~ is proportional to (I) differentiated by time. Using the change of fluxoid due to AC magnetic field, the magnetic penetration depth ( 2 ') is calculated as
1] '=R[i-[1 ~R"i aba~ "T
(1)
where R is the radius of the specimen. The relation between b and /l ' is obtained from eq.(1). The slope between b and /l is concerned with J ~. According to the Bean' s critical state model, /l ' is approximately expressed as (2) )
~
/loJc The slope of b - / l represents 1 / / . t 0 J ~. This method of obtaining J ~ is called Campbell method. In the measurement of Y-Ba-Cu-O by the AC method, two slopes are observed; one is concerned with the transport critical current density, the other is concerned with the local critical current density. Using these two, J c is investigated.
MEASUREMENT AND DISCUSSION The 20 wt.% Ag doped rod-shaped YBCO superconductor which was made by the zone melting process was used as the sample [3]. The sample size is a radius of 2.215mm, length of 12mm. A current( I ) of 1~ [kHz] from 10 to 140 [mA] were supplied into the coil and then b from 9.59 • 10-4 to 1.34 • 10-2 [T] were generated. At the same time, DC magnetic flux density ( B ) from 0 to 0.03 [T] is supplied parallel to the specimen. The specimen is maintained at a temperature of 77.3 [K]. In the SRPM method, /l is obtained from the map of A R and A X as described above. At B - 0 . 0 1 and 0.02 [T], on account of the supplied DC magnetic field, 2 appears at b - 0. Taking it into consideration, /l ' of the SRPM method is shown as Fig.3. It is apparent that the slope between b and /l ' is linear especially when B is 0.01 and 0.02 [T]. It had been expected that the two slopes were available also in the SRPM method, but in this result the only one slope was obtained. Fig.4 shows the profile obtained in the AC method. Compared with b - 2 ' of the SRPM method, the value of /l ' at the region of strong b is quite similar, but in the SRPM method the value at B - 0 is different from the others. In the SRPM method, We tried to calculate the critical current density using the inclination
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Fig.2 Map of A R - A X
Fig.4
b-
703
Fig.3 b - /1 ' of SRPM method
2 ' of A C m e t h o d
Fig.5
J ~ in SRPM and AC methods
of the slope. J ~ obtained in the SRPM and in the AC method is shown in Fig.5. In the SRPM method, the result of J ~ under the DC magnetic field of 0 ~ 0.03 IT] was obtained. J ~ obtained in the SRPM method is similar to that of the AC method. J ~ obtained by the VSM is 1.7 • 10 ~ [A/m ~ ], and by the DC 4-probe method is more than 1.6 • 10~ [A/m ~ ] (as shown in table 1). All of the method could have nearly the same value. However, the appropriate calibration concerned with the effect of the shape of the specimen and the existence of cracks and normal conductive material in the superconductor on the value of 2 could not have been done in the SRPM method. Table1. Comparison of J ~ ( at 77.3 [K] ) B [Wb/m x]
0.01
O.
02
SRPM method
AC
method
1.039 • 10
3.063 • 10 ~
9.549 • 10
1.990 x 10
7.520 • 10 7
1.816 •
10
VSM
1.7 •
10
4-probe method
> 1.6 • 10 ~
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CONCLUSION When J c of rod-shaped sample was measured by the SRPM and AC methods under AC and DC magnetic flux density, J ~ obtained in both methods agrees well. It is obvious that there is a possibility to be able to apply this method to the measurement of critical current density. In the AC method, two slopes concerned with the transport critical current density and with the local critical current density are observed, but in the SRPM they are not observed. We have to investigate the reason. And moreover, the appropriate calibration concerned with the effect of the shape of the specimen and the existence of cracks and normal conductive material in superconductor on the value of /l ought to be done in the SRPM method. REFERENCES H.Nakane et.al, Calculation of The Difference in Impedance for a Solenoid Coil with and without a Sample Conductor, IEEE Trans. Instrum. Meas., (1991) 40 544-548 E. S. Otabe et.al, Estimation of Critical Current Density in a Melt-Processed Superconducting Y-Ba-Cu-O Using AC and DC Inductive Methods, Jpn. J. Appl. phys. (1994) 33 996-999 S. Nagaya et.al, Ag Doping Effects on the Microstructure and Properties of Unidirectional Grown Y-Ba-Cu-O Superconductors, IEEE Trans. Magn. (1995) 5 1564-1567
AUTHOR INDEX
Abe, T., 1907 Abe, Y., 1333 Agapov, N., 139 Agatsuma, K., 1701 Ageta, T., 891,899 Aihara, K., 1261, 1685, 1689 Aiyama, Y., 891 Ajima, Y., 419 Akazaki, T., 1193 Akhmetov, A., 1829, 1833 Akinaga, M., 1647 Akita, S., 933 Aksenova, E. N., 1937 Akune, T., 1443, 1561 Alexeev, A., 395 Amano, K., 1707 Amardas, A., 819 Amemiya, N., 1617, 1775 Amrit, J., 559 An, C.-W., 847, 1269 Ando, M., 493 Ando, T., 763, 767, 771,775, 783,787, 791, 795, 807, 1265, 1661, 1665, 1673, 1677, 1895, 2009 Andoh, H., 1571 Andrikidis, C., 1353, 1545, 1549 Antipov, E. V., 1627 Aoki, K., 867 Aoki, N., 1681 Aoki, R., 399 Arai, K., 1297, 1701 Araki, S., 747 Araki, T., 263, 307 Arasawa, K., 267 Arata, M., 1273 Arend, I., 523, 547, 693 Asakura, H., 203, 207 Asakura, S., 999 Asami, H., 351 Asano, K., 751, 755, 759, 811, 1017, 1049, 1989, 2035, 2061 Asano, T., 1089 Aubin, M., 1621 Awaji, S., 1109, 1121, 1365, 1669, 1719, 1723, 1841 Ayai, N., 1673, 1677 Aymar, R., 53 Azuma, N., 1779 Baba, T., 75, 79, 83, 87, 735 Babic, E., 1557
Baenitz, M., 1561, 1627 Baixeras, J., 1479 Ballarino, A., 1139, 1143, 1147 Ban, T., 1655 Barclay, J. A., 2065 Barranco-Luque, M., 103 Batin, V. I., 439 Baudouy, B., 563 B/iuerle, D., 1533 Beales, T. P., 1541 Beduz, C., 1541 Benda, V., 199 Bender, S., 311 B6zaguet, A., 91, 519 Bhattacharya, D. G., 643 Bhattacharya, J. L., 949 Bhattacharya, N. C., 843 Bi, Y.-F, 151,847 Bian, S.-X., 287 Bianco, M., 1499 Binneberg, A., 505, 509 Biswas, B., 823, 827 Blackburn, T. R., 1037 Blau, B., 1665 Bocquillon, A., 435 Boehm, J., 1117 Bohno, T., 1053, 1151 Boiko, B. B., 1949, 1953 Bon Mardion, G., 689 Bona, M., 1911 Bondarenko, S. I., 1177 B6rner, H., 669 Bose, T. K., 2065 Boucheffa, A., 559, 563 Bourque, R., 427 Brandst/itter, G., 1579 Brehm, H., 259 Breitzke, H., 1627 Bremer, J., 173 Brissette, Y., 1021 Brunovsk2), I., 195 Bruzek, C. E., 1321 Bruzzone, P., 1243 Buhler, S., 449 Bunce, G., 871,875 Burger, J. F., 391 Buschmann, H., 509 Cai, J.-H., 271 Cai, X., 1189 Calvone, F., 1871
Camille, R. J., 803 Casas-Cubillos, J., 91 Cave, J. R., 1021, 1621 Chaffard, P., 107 Chahine, R., 2065 Chandratilleke, R., 501 Chauhan, H. S., 1509 Chaussonnet, P., 435 Chen, C.-Z., 217, 221, 225, 229, 233, 497 Chen, G.-B., 247, 275, 315, 407, 639, 677 Chen, G.-M., 407, 639 Chen, H. Y., 385, 677 Chen, Y., 359 Cheng, S.-K., 1447 Chensky, I., 799 Chiba, A., 1907 Chiba, H., 1847, 1851, 1863 Chida, K., 87 Chiba, M., 1711, 1847 Chien, S. B., 385 Chikaba, J., 1369 Chikaraishi, H., 735, 739, 743, 747, 751 Choi, I.-H., 1467 Choi, K.-D., 941,953 Chubraeva, L. I., 895, 925, 929, 1933, 1937 Claudet, S., 103 Collaudin, B., 711 Collings, E. W., 1397, 1609, 1767 Cowey, L., 1143 Cragg, D., 111, 115, 123 Cruikshank, P., 681 Cui, G.-W., 677 Cullen, J. R., 871,875 Cur6, C., 111, 123 Daido, K., 481 Daniels, P., 1117 Das, S., 823, 827 Das, S. K., 643 Dauvergne, J. P., 107, 111, 169, 173, 199 Decker, L., 195 Delikaris, D., 169, 173 Delruelle, N., 173 De y, R., 643 Dhard, C. P., 819 Din, L. R., 847
12
A u t h o r Index
Doi, Y., 119, 165, 419, 843, 867 Doko, T., 1915 Dolez, P., 1621 Dolgosheev, P., 799 Dong, D., 719 Dou, S. X., 1037, 1353, 1393, 1397, 1401, 1545, 1549, 1557, 1609 Drozd, A. A., 1941, 1945, 1953 Duan, Z., 1373 Durga Prasad, K. A., 949 Duthil, R., 123 Ebihara, K., 1483 Egawa, K., 1731, 1771 Egi, T., 1463, 1467 Emoto, M., 673, 1285 Endo, A., 1509 Endo, S., 182 I Enoki, T., 1193 Erbuke, L., 863 Eriksson, H., 863 Ernesto, M., 983 Escher, U., 2023 Ettlinger, E., 707 Evans, D., 2017 Evans, L. R., 45 EXSIV Group, 739, 743, 747 Ezaki, T., 1045, 1049 Fabian, P. E., 1997 Ferlin, G., 443 Filippov, Yu. P., 439, 609 Fisk, H. E., 863 Fleshier, S., 1613 Flokstra, J., 1181 Flukiger, R., 863 F61de~iki, M., 2065 Franco, C., 983 Frangois, M. X., 559, 563 Frederking, T. H. K., 657 Friend, C. M., 1541 Fuchino, S., 593, 957, 1297 Fuji, H., 1013, 1867 Fujii, M., 473 Fujii, Y., 263, 461,567, 621, 1525 Fujikami, J., 967, 975, 1347, 1597 Fujima, K., 1297 Fujimoto, H., 1655 Fujimoto, S., 331, 1173 Fujinami, T., 331 Fujino, K., 1413, 1583 Fujinuma, S., 1073 Fujioka, K., 291, 513, 2069 Fujioka, T., 767, 783 Fujisaki, H., 779 Fujishima, S., 855, 859 Fujita, T., 183, 665 Fujiyama, N., 381 Fujiyoshi, T., 1505 Fukano, T., 87 Fukasaku, Y., 279 Fukuda, K., 1669, 1685 Fukuda, T., 2069, 2073 Fukuda, Y., 1693 Fukui, N., 1173 Fukui, S., 1775
Fukushima, K., 1521 Fukuyama, S., 1919 Funaki, K., 1009, 1049, 1053, 1325, 1329 Funayama, H., 1727 Furumoto, H., 403, 601 Furuto, Y., 913, 917, 921 Furuya, S., 267 Futaki, N., 1381, 1601 Ganzinov, 1. S., 929 Gao, J. L., 295, 303 Gao, C.-H., 221,497 Gauthier, A., 519 Gayet, Ph., 103 Geilenkeuser, R., 2031 Genoud, J.-Y., 1571 Gerstenberg, H., 2001 Gerster, J., 259 Geynisman, M. G., 143 Ghose, D., 643 Ghosh, A. K., 863 Ghosh, G., 1957 Giesey, R. K., 2053 Gifford, P. E., 363 Gippius, A. A., 1627 Gistau-Bagucr, G., 189 Gladun, A., 2023 Godeke, A., 803 Gong, L.-H., 343, 347 Goodrich, L. F., 1715 Gopal, R. B., 2065 Goto, A., 855, 859 Goto, K., 933, 1073, 1197, 1723, 1841, 1867 Goto, T., 1189 Gradt, T., 669 Grantham, C., 1037 Gravil, B., 689 Green, M. A., 871,875 Gregory, E., 1715 Grimaud, L., 519 Grman, D., 863 Grunblatt, G., 1321 Gu, A.-Z., 581 Guinaudeau, H., 91 Gulko, E., 1715 Gtintherodt, H.-J., 1627 Guo, Y.-Y., 287, 613 Guo, Y. C., 1037, 1353, 1393, 1401 Haberstroh, Ch., 395 Hahakura, S., 1347 Hahn, S.-Y., 941 Hakamata, M., 1775, 1847 Hakuraku, Y., 1491, 1495 Hama, K., 535 Hamada, H., 1525 Hamada, K., 127, 131, 135, 427, 493, 767, 783 Hamada, M., 1061, 1095, 1293 Hamada, T., 1443 Hamaguchi, S., 539 Hamasaki, K., 1197 Hamashima, T., 1727 Han, S., 1513
Han, G., 1919 Hanaoka, Y. W., 1685, 1689 Hangyo, M., 1205, 1209, 1631 Hara, K., 183, 851, 1249 Hara, M., 1009 Hara, N., 1305, 1309, 1313 Hara, T., 963, 967, 975, 979, 1029, 1033, 1413, 1575, 1583, 1597, 1605 Harada, N., 1693 Harada, S., 1583 Haraguchi, E., 1173 Harrison, S., 1143 Hartwig, G., 1977 Haruyama, T., 119, 419, 513,649, 843, 867 Hasanain, S. K., 1517 Hase, N., ! 571 Hase, T., 1409 Hasebe, T., 1109, 1121 Hasegawa, H., 909 Hasegawa, K., 1413 Hasegawa, M., 771 Hasegawa, T., 1361 Hasegawa, Y., 399 Hashimoto, T., 325, 1377 Hata, K., 535, 585, 617 Hatakeyama, H., 501, 1113 Haug, F., 107, 111, 123, 169, 173, 199 Hayakawa, N., 1837 Hayashi, H., 735, 1017, 1151 Hayashi, K., 1009, 1329, 1347, 1413 Hayashi, S., 1129, 1409 He, J.-H., 1919 Heiden, C., 283, 311,453 Heinze, M., 1627 Herz, W., 469 Herzog, H., 147 Herzog, R., 505, 509 Hideto, Y., 335 Higashi, N., 419, 851, 1249 Higo, S., 1491, 1495 Higuchi, K., 1103 Higuchi, N., 1297 Hikata, T., 1347 Hikita, M., 1837 Hilbert, B., 91 Hino, N., 1305 Hirabayashi, H., 513, 851 Hirai, H., 2009 Hirakawa, K., 1451 Hirano, N., 1227 Hirano, S., 1583, 1597 Hirao, Y., 211 Hirayama, T., 1455 Hiresaki, Y., 303 Hirose, R., 1095, 1293 Hirumachi, T., 783 Hisada, S., 665 Hishinuma, Y., 1427 Hiue, H., 135, 743, 747 Hiyama, T., 127, 493 Hoffmann, W., 1627 Hofmann, A., 239 Hojo, M., 1791 Hommei, T., 1309, 1313
Author Index Honda, T., 127, 131, 135, 427, 493, 767, 787, 795 Honda, Y., 1277 Honjo, S., 963, 975, 979, 1575, 1605 Honjyo, S., 967, 1413 Honma, H., 1073 Honmei, T., 1305 Horigami, O., 1361 Horiguchi, K., 1887, 1981 Horise, R., 1129 Horiuchi, T., 1103 Horiya, T., 1895, 1915, 1923 Horvat, J., 1353, 1401, 1557 Horvath, I. L., 863 Hosaka, M., 1003 Hoshino, T., 941,953, 1639 Hosokawa, M., 1041 Hosoya, T., 203 Hosoyama, K., 183, 851, 1239, 1249 Hotta, Y., 811 Huang, B. J., 385 Huang, J.-H., 581 Huang, X., 571 Huang, Z.-P., 359 Huang, Z.-X., 247, 275, 677 Hubler, U., 1627 Htibner, R., 1977 Htibner, W., 669 Humer, K., 1997, 2005 Hussain, M., 2057 Huwiler, R., 863 Ichige, K., 493 Ichihara, M., 1681 Ichihara, T., 771 lchikawa, T., 905, 991, 1847 Ichikohara, H., 1693 Ichimaru, O., 1817 Ichiyanagi, N., 979 lde, Y., 1561 li, H., 1405, 1575 Iida, F., 135 Iida, M., 419 Iida, T., 1915 Iijima, Y., 995, 1003, 1669, 1697 lima, M., 735 Iimura, K., 79 Ijspeert, A., 1139, 1143, 1147 lkeda, H., 1385, 1435 Ikeda, Y., 159, 211 Ikegami, K., 855, 859 Ikegami, T., 1483 Ikeuchi, M., 159 Ikeya, T., 2069 Ikuhara, Y., 1455 Imagawa, S., 735, 739, 751,755, 759, 1825 Imai, Y., 909 Imayoshi, T., 1017, 1049, 1151, 1829 Imokawa, H., 1201 Inaba, S., 751 Inagaki, J., 783 Inoue, K., 913, 917, 921, 1089, 1099, 1103, 1361, 1409, 1669, 1697, 1735 lnoue, N., 735 Inoue, T., 299
Inoue, Y., 1129, 1711 Ioka, S., 735 Ionescu, M., 1353 Ipatov, Y., 1969 Irie, F., 1049, 1151, 1567 Ishibashi, K., 1163 Ishige, K., 1923 lshigohka, T., 513, 1069, 1337 lshihara, M., 1109, I 121 lshii, H., 621,963,967, 975, 979, 1413, 1575, 1583, 1597, 1605 Ishii, I., 593, 957, 1297 lshikawa, J., 831 Ishiyama, A., 513, 1239, 1281 lshizuka, M., 1377, 1385 lsogami, H., 1125 Isojima, S., 617, 967, 971, 975, 1057, 1347 Isono, T., 767, 771, 775, 1265, 1731, 1771 Ito, D., 1553, 1821 Ito, M., 1347, 1775 Ito, T., 779, 791, 1301 Itoh, A., 427, 1643 Itoh, H., 685 Itoh, I., 743, 795, 1325, 1329 Itoh, K., 1103, 1735, 1783, 1787, 1799 Itoh, M., 1499 Itoh, R., 1049 Itoh, S., 431, 1129 Itou, I., 807 Ivano, O., 1895 Iwabuchi, A., 1727, 1907, 1961, 1965 Iwaki, G., 1669, 1685, 1689, 1719 Iwakuma, M., 1009, 1325, 1329 lwamoto, A., 75, 79, 83, 87, 605, 735, 739, 743, 751, 1227, 1253 Iwamoto, K., 71, 155 Iwamoto, S., 807 iwasaki, S., 1723, 1841 lwasaki, T., 203 Iwasaki, Y., 2005 Iwaski, S., 933, 1443 J/ickel, M., 2023, 2031 Jacob, S., 413, 949 Jaffery, T. S., 1235 Jager, B., 435, 689 Jayakumar, J., 763 Jenninger, B., 443 Jeong, S.-K., 1215, 1219, 1223, 1231 Jess, P., 1627 Ji, P., 1447 Jikihara, K., 1109 Jin, J. X., 1037 Jochimsen, G., 423, 707, 715 Juillet, J. J., 711 Kabashima, S., 539 Kabe, A., 183, 851, 1249 Kai, T., 739, 743 Kaiho, K., 1013, 1297 Kaiser, G., 259 Kaito, T., 1057 Kajikawa, K., 1013, 1297 Kakehi, Y., 1201, 1591
13
Kakimi, Y., 291 Kakugawa, S., 1305 Kalinin, V., 131 Kamata, K., 1851, 1863 Kamada, S., 1431 Kamikado, T., 1293 Kamioka, Y., 457 Kamiya, I., 1045 Kamiya, K., 531 Kanari, T., 543 Kanazawa, Y., 339, 351,355 Kanda, Y., 1337 Kanegae, K., 1329 Kanekiyo, T., 183 Kaneko, T., 1347 Kancko, Y., 203, 207 Kang, Y.-M., 629, 1173 Karimoto, S., 1487 Kariya, J., 673 Karunanithi, R. 413, 949 Kasagawa, Y., 1009, 1325, 1329 Kasahara, H., 933 Kasahara, S., 331 Kashima, T., 1993, 2049 Kasthurirengan, S., 413, 949 Kasuga, T., 1333 Kasuu, O., 1155 Kasuya, M., 267 Katada, M., 67, 83 Katheder, H., 1665, 2001 Kato, H., 1929 Kato, S., 419, 843 Kato, T., 127, 131, 135, 427, 493, 767, 771,775, 795, 1347 Katoh, Y., 735 Katsumura, Y., 1357 Kauschke, M., 465 Kawabata, C., 1651 Kawabata, S., 921 Kawaguchi, E., 263, 307 Kawaguchi, T., 855, 859 Kawai, M., 419, 867 Kawakami, A., 1185 Kawamata, H., 419, 843, 851, 1249 Kawakami, K., 551 Kawano, H., 1305, 1309, 1313 Kawano, I., 1491 Kawano, K., 127, 473, 493 Kawano, S., 299 Kawasaki, K., 1435 Kawashima, I., 127 Kawate, Y., 431,513, 1103, 1129, 1409 Kazumori, M., 885 Kesseler, G., 173 Khodzhibagiyan, H., 139 Kiboshi, T., 1753 Kida, J., 1103 Kido, G., 1089 Kido, T., 1169, 1173 Kikuchi, A., 1471 Kikuchi, K., 1685, 1689 Kikuchi, M., 1681 Kim, J.-W., 855, 859 Kim, S., 419, 843 Kim, S.-K., 1281 Kim, S. W., 1767
I4
Author Index
Kim, Tae Hyun, 1133 Kimura, A., 913, 917, 921, 1575, 1605 Kimura, H., 1073, 2069, 2073 Kimura, M., 1669, 1719 Kimura, N., 419, 649, 843, 1129 Kimura, Y., 851 Kirby, G., 837 Kishida, T., 1057 Kisida, T., 909 Kiss, T., 1567, 1587 Kitagawa, K., 211 Kitaguchi, H., 1099, 1957 Kitamura, M., !305 Kitamura, R., 1803 Kiyoshi, T., 431, 1089, 1099, 1103, 1361, 1409, 1735 Klundt, K., 283 Knoopers, H. G., 803 Knoops, S., 199 Koba, S., 1491, 1495 Kobayashi, H., 551 Kobayashi, N., 1669, 1707, 1719, 1739 Kobayashi, S., 1347 Kobayashi, T., 863, 867, 1189 Kobori, T., 1887 Kodama, T., 307 Koga, T., 1069 Kohler, C., 1321 Kohno, O., 933, 987, 995, 999, 1003, 1013, 1073, 1381, 1601, 1723, 1841, 1867 Koike, T., 1083 Koizumi, N., 767, 771, 775, 791, 795, 1301, 1673, 1677 Koizumi, T., 1361 Kojima, Y., 183, 851, 1249 Komatsu, K., 885 Komatsu, M., 1537 Konda, H., 1993 Kondo, T., 119 Kondo, Y., 867 Kondou, Y., 419 Konno, M., 795, 1009, 1053, 1151, 1325, 1329, 1665 Konosu, S., 1895 Kos, N., 681 Kosaka, T., 381 Koshizuka, N., 1463, 1467 Kosuge, M., 1099, 1697 Kosugi, K., 1405 Koto, S., 701 K0uki, N., 335 Kouriki, K., 811 Kovalenko, A., 139 Kovrizhnykh, A. M., 439 Koyanagi, K., 1113, 1707 Kr/ihling, E., 2001 Kreisler, A., 1479 Krempetz, K., 863 Krooshoop, H. J. G., 803 Kubo, S., 1487 Kubo, T., 855, 859 Kubo, Y., 1731, 1771, 1799 Kubota, H., 1749, 1763 Kubota, Y., 1763, 1779 Kuchiishi, Y., 751
Kukano, T., 83 Kuma, S., 1389 Kumakura, H., 1099 Kumano, T., 1799 Kume, A., 987, 1381, 1601 Kundzins, K., 1579 Kurahashi, H., 1739 Kurihara, T., 331, 367 Kuriyaki, H., 1451 Kuriyama, F., 279 Kuroda, K., 1463, 1467, 1833 Kurtyka, T., 1911 Kurusu, T., 1029, 1201 Kusayanagi, E., 1817 Kusevic, I., 1557 Kushida, T., 1277
Maki, N., 1305, 1309, 1313 Makida, Y., 119, 419, 867 Mamalis, A. G., 937 Manzoor, S., 1517 Mao, C., 1421 Mao, C. B., 1417 Mao, D., 1783 Marechal, J. L., 689 Marque, S., 1911 Marti, H. P., 863 Martinez, A., 435 Maruno, Y., 597 Masada, E., 1041, 1079 Masashi, N., 335 Masatomi, H., 399 Masegi, T., 1681, 1707 Masuda, T., 971, 1057, 1347 Landgral', R., 283 Masulnoto, T., 1121 Lang, H. P., 1627 Masuzaki, S., 735 Le Lay, L., 1541 Matsubara, Y., 291,295, 303, 319 Lebrun, Ph., 91, 95, 195, 199, 443 Matsuda, H., 67, 71, 83 Lehmann, W., 489 Matsui, K., 127, 131, 135, 493, 767, LHD Group, 63, 75, 79, 83, 87, 731, 771,775, 779, 795, 1301, 1681 751, 1825 Matsui, T., 299 Li, B. Z., 847 Matsukawa, M., 1073, 1841 Li, J. N., 1037 Matsukura, N., 1711 Li, L.-Z., 1269 Matsumoto, K., 1095, 1129 Li, R., 339, 351, 355 Matsuo, M., 1053 Li, S., 359 Matsuo, S., 403 Li, X. Y., 1037 Matsuoka, S., 1459 Li, Y.-Y., 677 Matsushita, T., 701, 1795, 1803 Li, Y.-Z., 287, 497, 523, 547, 693 Matsuzawa, H., 831 Li, Z., 581 Mawatari, Y., 1529 Li, Z.-M., 461 Mayaux, C., 485 Li, Z.-Z., 613 Mayri, C., 111, 123 Liang, J.-T., 271 Maytal, B.-Z., 635 Lierl, H., 147 Mazaki, H., 1537 Lin, L. Z., 847, 1513 Mazurenko, O. N., 1941, 1945, 1949, Lin, X.-J., 871,875 1953 Liu, H. K., 1037, 1353, 1393, 1397, McIntyre, P., 1235 1545, 1549, 1557 Melaaen, E., 99 Liu, H. L., 1037 Meslmani, Y., 1533, 1579 Liu, J. Y., 1037 Meuris, C., 563 Liu, L.-Q., 225, 229 Michael, P., 1215, 1219, 1223 Liu, R.-M., 677 Michael, P. C., 1231 Liu, X.-Y., 2065 Mikawa, M., 1197 Liu, Z. Y., 1037 Mikumo, A., 1661, 1673, 1677 L6hlein, K., 195 Miller, J. R., 1891 Lokken, O. D., 363 Mimori, K., 419 Lounasmaa, O. V., 27 Mimura, M., 979, 1405, 1575, 1605 Lu, X. Y., 1431 Minami, H., 1083 Luciano, M., 983 Minami, M., 1961, 1965 Lfiders, K., 523, 547, 693, 1561, 1627 Minato, T., 771 Luo, E., 271 Mine, S., 863 Lutset, M., 1341 Minemoto, T., 1499 Minervini, J., 763, 1215, 1219, 1223 Machi, T., 1463 Minervini, J. V., 803 Machida, A., 159 Minot, F., 689 Maeda, M., 461, 621 Misaki, Y., 1197 Maeda, H., 431, 1089, 1099, 1103, Mitchell, N., 763 1377, 1385, 1735, 1757 Mitchell, N. A., 1903 Maehata, K., 1163 Mitin, V., 653 Maekawa, R., 75, 79, 83, 87, 481,735, Mito, T., 63, 75, 79, 83, 87, 605, 735, 739, 743, 1227 739, 743, 747, 751, 1163, 1227, 1253, Maezono, K., 1495 1337, 1825 Maix, R., 2001 Mitrohin, V., 799
Author Index Mitsubori, H., 1109 Mitsui, H., 1273, 2009 Mitsumoto, T., 855, 859 Miura, A., 331 Miura, K., 331,367 Miura, O., 1821 Miura, Y., 775, 1265, 1301 Miyaike, K., 905 Miyaji, T., 751 Miyake, A., 131,203, 743 Miyashita, K., 1851, 1863 Miyatake, T., 1095, 1711, 1735, 1739 Miyauchi, Y., 427 Miyazaki, T., 1095, 1711, 1739 Miyoshi, K., 979 Mizumaki, S., 419 Mizusaki, K., 1443 Mizutani, Y., 831 Moca6r, P., 1321 Mogi, I., 1121 Mohr, D., 423 Momal, F., 199 Mori, K., 1305, 1499 Mori, M., 131,203 Mori, S., 457 Moriai, H., 751 Morimoto, H., 267, 1017 Morisaki, T., 735 Morishita, H., 367 Morita, H., 811, 1261 Morita, M., 1125, 1289, 1731, 1771 Morita, Y., 183, 851, 1249 Moriuchi, S., 75, 79, 83, 87, 735 Moriya, T., 1763 Moriyama, H., 1273, 2009 Morra, M. M., 1903 Motojima, O., 63, 67, 75, 79, 83, 87, 673, 725, 735, 739, 743, 747, 751, 1159, 1285, 1825 Motokawa, M., 1121 Mtick, M., 283 Mukai, E., 941,953 Mukai, H., 1413 Mukoyama, S., 979 Mfiller, K.-H., 1353, 1545, 1549 Munshi, N. A., 1997 Murai, K., 79 Murai, S., 783 Murakami, M., 251, 255, 531, 543, 575, 589, 1439 Murakami, T., 2045 Murakami, Y., 1041, 1205, 1209, 1631, 1783 Murase, S., 1681, 1707, 1795 Murata, Y., 811 Muta, I., 941,953 Nabatame, T., 1685, 1689 Nadi, R., 1021, 1621 Nagai, K., 1957 Nagai, T., 1731, 1771 Nagamura, H., 909 Nagano, M., 1537 Nagata, A., 1431 Nagata, M., 987, 1749
Nagaya, S., 701, 971, 987, 995, 999, 1381, 1601 Nakade, M., 1029, 1033 Nakagawa, H., 1923 Nakagawa, M., 987, 999, 1381, 1601, 1701 Nakagome, H., 319, 501, 1033, 1113 Nakahara, S., 665 Nakahira, A., 2041, 2057 Nakahira, M., 427 Nakai, H., 183, 851, 1249 Nakajima, H., 767, 771,775, 783, 795, 1665, 1887, 1895 Nakamoto, K., 739 Nakamoto, T., 419 Nakamoto, Y., 331 Nakamura, H., 913, 917, 921 Nakamura, K., 79, 83, 473, 1333 Nakamura, M., 263, 461, 567, 621, 1455, 1525 Nakamura, N., 251,255 Nakamura, T., 1567, 1587 Nakane, H., 701 Nakanishi, K., 755, 759 Nakaniwa, T., 1895 Nakao, H., 331 Nakayama, S., 885, 891, 1707 Nakayama, Y., 127 Nara, K., 1929 Narayankhedkar, K. G., 377, 373 Natori, N., 1297 Natu, P. V., 373 Nemoto, T., 1993 Nemoto, Y., 1757 Neo, S., 1041 Neubert, J., 509 Neuenschwander, J., 863 Nicoletti, A., 879 Nicollet, S., 435 Nii, A., 1025 Niihara, K., 2041, 2057 Nijhuis, A., 1243 Ninomiya, A., 735, 1069, 1337 Nishi, M., 1731, 1771 Nishida, K., 127, 131, 135, 427, 493, 767, 795 Nishigaki, K., 513, 597, 1813 Nishiguchi, K., 1173 Nishihara, R., 1505 Nishijima, S., 19, 513, 1273, 1277, 1989, 2013, 2035, 2041, 2049, 2057, 2061 Nishikawa, M., 891 Nishimura, A., 735, 739, 743,751,755, 759, 1877 Nishimura, I., 203 Nishimura, K., 735 Nishioka, T., 1163 Nishitani, T., 263, 307 Nishiura, T., 1989, 2013 Nishiya, T., 941,953 Nisiwaki, Y., 1065 Nitta, I., 1961, 1965 Nitta, J., 1193 Nitta, T., 891,909, 1065 Nobutoki, M., 71
I5
Nogawa, S., 1065 Noguchi, T., 513 Nojima, K., 2035, 2061 Noma, K., 2005 Nomura, H., 1065, 1297 Nomura, K., 1389 Nomura, S., 1029, 1113, 1707 Nonaka, S., 1357 Norris, B. L., 179 Nose, S., 1009, 1049, 1053, 1151, 1325, 1329 Noto, K., 1073, 1841 Nozaki, S., 1061 Nozawa, M., 791,795, 1265 Numazawa, T., 2069, 2073 Nunoya, Y., 767, 771, 775, 779, 795, 1665, 1731, 1771, 1895 Oba, K., 79 Obara, H., 1529 Obert, W., 485 Ochi, T., 399 Oellrich, L. R., 239 Ogasawara, M., 1471 Ogasawara, T., 1763, 1779 Ogata, H., 775, 783 Ogata, T., 1791, 1899, 1915, 1923 Ogawa, H., 735 Ogawa, R., 1103, 1129, 1409, 1711, 1799 Ogawa, S., 1201, 1591 Ogino, O., 513 Ogiso, K., 1041 Ogitsu, T., 183, 419, 843 Ogiwara, H., 513 Ogushi, T., 399, 1443, 1491, 1495 Ohashi, Y., 299 Ohba, K., 87, 735 Ohhata, H., 419 Ohira, K., 403, 601 Ohkita, S., 1895 Ohkuma, T., 1029, 1033, 1413 Ohkura, K., 1057, 1347, 1365 Ohmatsu, K., 967, 975, 1347, 1799, 1855, 1859 Ohno, I., 427, 743 Ohsaki, H., 1079 Ohsaki, O., 783 Ohska, T., 673 Ohtake, I., 87, 735 Ohtani, Y., 319, 501, 1113 Ohtsu, K., 493 Ohtsuka, H., 1957 Ohuchi, N., 419, 843 Okada, H., 1049 Okada, K., 473 Okada, T., 909, 1273, 1277, 1989, 2013, 2035, 2041, 2049, 2057, 2061 Okaguchi, S., 1915, 1923 Okaji, M., 1929 Okamoto, M., 331 Okamura, T., 539 Okano, M., 593, 957 Okazaki, O., 513 Okubo, H., 319, 1837 Okubo, K., 1009, 1325, 1329
I6
Author Index
Okumura, H., 673 Okuno, H., 855, 859 Okuno, K., 763, 791 Onabe, K., 995 Onishi, A., 339, 351,355 Onishi, T., 1025 Ono, M., 739, 987, 1727 Onoda, H., 1361 Onodera, T., 1841 Ooba, K., 75 Ootsu, K., 127 Ootuka, Y., 457 Osaki, O., 767 Osaki, K., 1693 Osamura, K., 1357, 1787, 1791, 1795, 1799 Oshikiri, M., 775, 1665 Ostler, J., 837 Otabe, E. S., 701, 1803 Otsuka, M., 791, 1985 Owren, G., 99 Ozaki, O., 1095, 1129, 1293 Ozeki, M., 13 Pal3vogel, Th., 707, 715 Pai, Chien-ih, 871 Pai, C., 875 Pailler, P., 111 Pan, H.-Y., 359 Panek, J., 571 Papavasiliou, N., 657 Passardi, G., 111, 123, 173 Passvogel, Th., 711 Patel, L. N., 377 Pavese, F., 1499 Peltier, F., 1321 Penny, M., 1541 Pe6n, G., 443, 477 Perini, D., 837 Peshkov, I., 799 Petersen, K., 423, 707 Pfotenhauer, J. M., 363 Plashkin, E. A., 1937 Pradhan, S., 819, 823, 827 Proyer, S., 1533 Pylinina, S. N., 929, 1933 Pyon, T., 1715 Qiao, G.-W., 1475 Qiu, L.-M., 247, 315 Qiu, M., 1513 Qiu, N., 151 Quack, H., 395, 465 Quan, H.-Y., 233 Radebaugh, R., 33 Raju, K. S. N., 949 Randall, R. N., 1903 Ravikumar, K. V., 657 Reed, R. P., 2017 Rehak, M. L. F., 879 Riddone, G., 95, 443, 477, 681 Rieder, H., 107 Rieubland, J. M., 173 Rodriguez Mateos, F., 1871 Rogalla, H., 391, 1181
Rohleder, I., 1665 Rousset, B., 519 Ruppert, U., 523, 547, 693, 945 Rychagov, A., 799, 1969 Ryouman, A., 1057 Sadakata, N., 933,987, 995,999, 1003, 1013, 1073, 1381, 1601, 1701, 1723, 1841, 1867 Saga, N., 967, 975, 1347 Sagner, U., 707 Saho, N., 1125 Sahu, A. K., 819 Saito, A., 1197 Saito, K., 1807 Saito, M., !915, 1923 Saito, T., 995, 1003, 1013, 1073, 1189, 1723, 1867 Saitoh, T., 933, 987, 999, 1381, 1601, 1701, 1795 Saji, N., 203, 207, 743 Sakagami, Y., 1681 Sakai, K., 1205, 1209, 1631 Sakai, S., 1669, 1685, 1689, 1693, 1719, 1791, 1799, 1851, 1863 Sakai, Y., 1089 Sakaki, K., 135, 795, 1053, 1151 Sakakibara, S., 735 Sakamoto, H., 913, 917, 921 Sakamoto, N., 1443, 1561 Sakamoto, Y., 183 Sakiyama, H., 665 Sakuma, S., 539 Sakuraba, J., 1109, 1121 Sakurai, A., 535, 585, 617 Salunin, N. I., 1937 Samadi Hosseinali, G., 1533, 1579 Samoto, K., 1369 Sampson, W. B., 863 Sanada, K., 1887, 1981 Sander, M., 423, 707, 715 Sang, I.-Y., 1439 Sapozhnicov, V. A., 929 Sarkar, B., 819 Sarrhini, O., 1479 Sasaki, K., 843, 1685 Sasaki, T., 771,783, 2009 Sashida, T., 1347 Sata, K., 1173 Satisha, G. V., 413 Sato, A., 431, 513, 1103, 2073 Sato, J., 203, 1389 Sato, K., 617, 913, 917, 921,967, 975, 991, 1009, 1057, 1155, 1329, 1347, 1365, 1413, 1583, 1597, 1673, 1677, 1855, 1859 Sato, M., 2069, 2073 Sato, S., 1017, 1053 Sato, T., 1041 Satoh, S., 63, 67, 75, 79, 83, 87, 481, 735, 739, 743 Satoh, T., 339, 351,355 Satoh, Y., 1049 Satou, K., 1405 Satow, T., 735, 739, 743, 747, 751, 1825
Sauerzopf, F. M., 1579 Sawa, A., 1529 Sawa, F., 1273, 2057 Saxena, Y. C., 819, 823, 827 Scanlan, R., 1235 Scanlan, R. M., 1743, 1767 Schauer, F., 815 Schultz, J., 1219, 1223 Schultz, J. H., 1231 Schultzand, J., 1215 Schumann, B., 505 Schupp, J., 715 Schustr, P., 195 Seeber, B., 863 Segawa, T., 1073 Seidel, A., 423, 707, 711,715 Seidel, P., 259 Seidler, M., 489 Seido, M., 751 Seki, N., 267 Sekiguchi, H., 75, 79, 87, 735 Sekiguchi, S., 127, 493 Sekine, S., 1297 Sekino, T., 2041 Semeonov, I., 799 Senba, A., 1079 Senba, T., 751, 1285 Seo, K., 251,255, 1289 Seppala, J., 863 Sergeyev, I. A., 609 Sergo, V., 199 Serio, L., 91 Serries, J. P., 435 Sgobba, S., 1911 Shamoto, Y., 673 Shavit, A., 635 Shen, M., 677 Shen, S., 1231, 1665 Shevchenko, O. A., 803 Shi, W., 2027 Shibata, K., 1883 Shibata, T., 967, 975, 1347 Shibutani, K., 1129, 1409 Shibuya, J., 775 Shieh, T. F., 385 Shiga, N., 1707 Shigematsu, T., 461, 567, 621, 1525 Shigenaka, A., 811 Shigi, T., 263, 461, 567, 621, 1525 Shimada, M., 1095, 1129, 1409, 1711, 1735, 1739, 1791 Shimakage, H., 1643 Shimamoto, S., 127, 131, 135, 427, 763, 767, 771, 775, 779, 787, 791, 807, 1301, 1665, 1673, 1677 Shimamura, K., 1707, 2073 Shimazaki, T., 575 Shimizu, K., 1069 Shimizu, T., 1727, 1907 Shimonosono, T., 971, 987, 995, 999, 1381, 1601, 1883 Shir.do, Y., 1887, 1981 Shingai, K., 1483 Shinohara, H., 1329 Shintomi, T., 419, 843,851, 1249, 1767 Shiohara, Y., 1455, 1459, 1509
Author Index Shioiri, T., 1239 Shiotsu, M., 535, 585, 617 Shiraishi, M., 251,255 Shoji, M., 673 Siegel, N., 837 Siemko, A., 837 Sigaev, V. E., 929, 1933, 1937 Simamoto, S., 1265 Simon, N. J., 2017 Singo, S., 1583 Sinha, B., 643 Sirot, E., 1321 Sirotko, D. V., 925 Skoczen, B., 1911 Smirnov, A., 139 Smith, B. A., 803 Smith, K., 1143 Smith, R. P., 863 Snydstrup, L. P., 871,875 Sobol, V. R., 1941, 1945, 1949, 1953 Solheim, N., 103 $611, M., 2001 Sotojima, T., 381 Specking, W., 1661 Spiebberger, S. M., 1997, 2005 Sp6rl, G., 509 SST Team, 819, 823, 827 Stamm, M., 489 Stangl, E., 1533 Starchl, B., 1579 Straif, W., 1579 Su, X.-D., 1475 Sfi/3er, M., 469 Sudo, S., 685 Suehiro, J., 1009 Suekane, T., 539 Sueyoshi, T., 1505 Suganomata, S., 831 Sugawara, K., 1431, 1753 Sugawara, S., 419 Sugimoto, M., 767, 771,775, 779, 783, 791,795, 913, 917, 921, 1605, 1665, 1673, 1677, 2009 Sugiura, T., 1571, 1841 Sugiyama, K., 1851, 1863 Sukhanova, A., 139 Sulten, P., 1321 Sumida, M., 1459 Sumita, T., 1749 Sumiyoshi, F., 921, 1049 Sumiyoshi, Y., 783, 2009 Summers, L. T., 1891 Sumption, M. D., 1609, 1767 Sun, T., 275 Sun, X. Y., 1417 Sunada, K., 629 Suraci, A., 91 Suryanarayana, T., 949 Suzawa, C., 617, 967, 975 Suzuki, H., 735 Suzuki, K., 905, 2045 Suzuki, M., 1451, 1487, 1639, 1753 Suzuki, N., 2045 Suzuki, S., 751,755, 759 Suzuki, T., 1239, 1685, 1689, 1985 Suzuki, Y., 1201
Svalov, G., 1969 Sytnikov, V., 799, 1969 Szalay, A., 937 Szeless, B., 443 Szeless, B., 1871 Szfics, Z., 945 Tachikawa, K., 1427, 1471, 1863 Tada, N., 1693 Taguchi, O., 1731, 1771 Taira, M., 1451, 1639 Takabatake, K., 1095, 1103, 1293 Takfics, S., 1253, 1257 Takagi, T., 1129 Takahashi, K., 1525 Takahashi, C., 1073 Takahashi, K., 1155, 1525, 1673, 1677, 1855, 1859 Takahashi, M., 1029, 1033 Takahashi, R., 811, 1261, 1317 Takahashi, T., 851, 1073 Takahashi, Y., 767, 771,775, 779, 783, 795, 811, 1301, 1665, 1681 Takahata, K., 79, 605, 735, 739, 743, 751, 1227, !253 Takano, K., 795, 1895 Takano, S., 1643 Takao, T., 1961, 1965, 1993 Takashi, I., 335 Takaya, Y., 1265 Takayanagi, H., 1193 Takayasu, M., 1215, 1219, 1223, 1231 Takebayashi, S., 1439 Takeda, M., 597, 1813 Takeo, M., 735, 1009, 1017, 1049, 1053, 1163, 1325, 1329, 1567, 1587, 1829, 1833 Takeshima, H., 1305, 1309, 1313 Takeshita, M., 621 Takeuchi, T., 1669, 1697, 1735, 1757 Takeuchi, Y., 535 Takigami, H., 913, 917, 921 Takita, K., 1159 Tallerico, T., 871,875 Tamada, N., 593, 957, 1297 Tamaki, T., 751,755, 759 Tamura, H., 735, 739, 755, 759 Tanahashi, S., 735, 739, 743 Tanahasi, S., 747 Tanaka, Y., 1385 Tanaka, A., 1361 Tanaka, H., 1833 Tanaka, K., 119, 419, 649, 843 Tanaka, M., 307 Tanaka, S., 3 Tanaka, T., 1883 Tanaka, Y., 855, 859, 979, 1377, 1385, 1405, 1575, 1605, 1787, 1799 Taneda, M., 127 Taneya, S., 331 Tang, H., 1475 Tang, X., 1373 Tang, Z.-M., 2027 Tani, M., 1205, 1209, 1631 Tanida, K., 303 Taniguchi, T., 1317
I7
Tanna, V., 819 Taran, A., 799 Tasaki, K., 1361 Tatara, I., 1739 Tateishi, H., 1297, 1701 Tavian, L., 95, 195, 199, 681 Tei, C., 1821 ten Haken, B., 697 ten Kate, H. H. J., 697, 803, 1243, 1743 Teng, M., 1143 ter Brake, H. J. M., 391, 1181 Terai, M., 331 Teramachi, Y., 673 Terasawa, A., 767, 771,791, 1265 Terashima, A., 419, 843, 851, 1249 Teuho, J., 863 Thome, R., 763 Thummes, G., 283, 311 Thfirk, M., 259 Timms, K., 1117 Tischhauser, J., 173 Titus, P. H., 1903 Tkhorik, Y., 653 Tobler, R. L., 1877 Tochihara, S., 1537 Toda, H., 1813 Togano, K., 1757 Tokunaga, M., 1749 Tominaga, A., 243 Tominaka, T., 855, 859, 1305 Tomioka, A., 1053, 1151 Tomioka, K., 331 Tommasini, D., 837 Tomozawa, S., 1205, 1631 Tonouchi, M., 1205, 1209, 1631, 1643 Torii, H., 367 Torii, S., 933 Toyoda, K., 905 Triscone, G., 1571 Troell, J., 453 Tschegg, E. K., 1997, 2005 Tsubouti, H., 979 Tsuchiya, K., 419, 843, 867 Tsugawa, K., 1297 Tsuji, H., 127, 131, 135, 427, 493, 763, 767, 771, 775, 779, 783, 787, 791, 795, 807, 811, 1265, 1301, 1661, 1665, 1673, 1677, 1731, 1771, 1887, 1895, 2009 Tsukamoto, H., 791, 1985 Tsukamoto, O., 1041, 1617, 1775 Tsukamoto, T., 1571 Tsukasaki, Y., 1989, 2013 Tsukiji, H., 941,953 Tsukiyama, M., 941,953 Tsuru, K., 1487 Tsutsumi, K., 1017, 1049, 1151 Tu~ek, L., 195 Tutaev, V. A., 929, 1933 Uchaikin, S. V., 439 Uchida, T., 2045 Uchikawa, F., 1731, 1771 Ueda, H., 1495 Ueda, N., 1083
I8
Author Index
Uede, T., 735, 743, 747, 1151 Uehara, M., 1957 Ueki, T., 2035, 2061 Ueno, S., 2041, 2057 Ueyama, M., 1009, 1329, 1347, 1365 Ulbricht, A., 469 Umeda, T., 203, 457, 1459 Uno, N., 979, 1405, 1575 Uno, S., 1057 Unoki, H., 1463, 1467 Urata, M., 1029, 1033, 1361, 1707 Uriu, Y., 1337 Usami, K., 1189 Usami, S., 1985 Ushigusa, K., 1681 Ushijima, I., 67, 83 Ushijima, M., 783 Utaka, Y., 1169 Uzawa, Y., 1185 Vajda, I., 937 Vallier, J. C., 435 Valthe, S., 2005 van Weelderen, R., 91, 519 van Oort, J. M., 1743 Van Sciver, S. W., 527, 571 Vanni, O., 983 Vanolo, M., 1499 Vecsey, G., 1665 Veldhuis, D., 1181 Venger, E., 653 Vieira, R., 2009 Vins, M., 195 Violet, J. L., 689 Vo, N. V., 1397, 1609 von Schoenebeck, F., 2023 Vullierme, B., 199 Vysotsky, V., 1215, 1219, 1223, 1231 Wachi, Y., 743 Wada, H., 431, 1089, 1099, 1103, 1783 Wada, K., 1427 Wada, N., 1209 Wadahl, A., 99 Wadayama, Y., 811, 1261 Wade, M., 1117 Wagner, A., 423 Wagner, R., 259 Wagner, U., 95, 99, 1139, 1143 Wakabayashi, H., 1301 Wakamoto, K., 1731, 1771 Wakasugi, K., 1749 Wakata, M., 1731, 1771 Wake, M., 867, 1235, 1767, 1807 Wakita, M., 1837 Wakuda, T., 1329 Walckiers, L., 837 Walker, E., 1571 Walker, R. J., 143 Walsh, R. P., 661, 1891 Wang, D., 677 Wang, F. T., 847 Wang, J.-R., 151 Wang, K.-G., 1447 Wang, L., 613 Wang, Q. L., 847, 1269
Wang, R.-S., 581 Wang, R.-Z., 555, 2027 Wang, W. G., 1353, 1393 Wang, W.-Y., 719 Wang, X.-X., 287 Wang, Y.-Q., 1269 Wang, Y.-Z., 1475 Wang, Z., 1185, 1205, 1209, 1631, 1643 Warnes, W., 1231 Watanabe, M., 975 Watanabe, I., 791 Watanabe, K., 79, 735, 1109, 1121, 1365, 1669, 1693, 1707, 1719, 1723, 1841 Watanabe, M., 457, 967, 975 Watanabe, N., 1045 Watanabe, Y., 127 Watazawa, K., 1109, 1121 Weber, J., 423 Weber, H. W., 1533, 1579, 1997, 2005 Weise, F., 2031 Welton, S. J., 527 Wen, J.-G., 1463 Wessel, S., 803 Wild, S., 239 Will6n, D. W. A., 1021, 1621 Williams, L. R., 477 Williamson, J., 657 Winkler, G., 103 Wolf, J., 423, 707 Wong, F. M. G., 1903 Wu, J.-Y., 555 Wu, J.-Y., 2027 Wu, X., 1373, 1421 Wu, X. Z., 693, 1417, 1447 Wu, Y., 217 Wu, Y.-Y., 547, 693 Wu, Y.-Z., 547, 693 Wtichner, F., 469 Wykes, M. E. P., 131,427 Xia, Z.-M., 247, 315 Xu, C., 1451 Xu, J.-M., 625, 1635 Xu, L., 271,625, 1635 Xu, R.-L., 677 Xu, X., 473 Xu, X.-D., 343, 347 Yabu-uchi, K., 399 Yaegashi, N., 1915, 1923 Yagi, N., 381 Yamada, H., 79, 87 Yalnada, N., 1929 Yamada, R., 863 Yamada, S., 63, 75, 79, 83, 87, 735, 739, 743, 747, 751 Yamada, Y., 1109, 1121, 1361, 1427, 1455, 1661, 1665, 1673, 1677 Yamafuji, K., 1009, 1325, 1329, 1567, 1587 Yamagata, Y., 1483 Yamaguchi, K., 905, 1017 Yamaguchi, M., 263, 461, 567, 589, 621, 1525
Yamaguchi, S., 673, 735, 751, 1159, 1285 Yamaguchi, T., 331 Yamamoto, A., 111, 119, 419, 649, 843, 867, 1915, 1923 Yamamoto, J., 63, 67, 75, 79, 83, 87, 605, 673, 731, 735, 739, 743, 747, 751, 755, 759, 1163, 1227, 1253, 1337, 1361, 1825, 1841, 1877 Yamamoto, K., 1029 Yamamoto, M., 263, 1013 Yamamoto, N., 1915 Yamamoto, S., 1133, 1289 Yamamoto, T., 743 Yamamoto, Y., 1525 Yamamura, H., 127 Yamanaka, A., 1993, 2049 Yamaoka, H., 419 Yamasaki, H., 1529 Yamasaki, S., 267, 1025 Yamashita, F., 1427 Yamazaki, K., 735 Yamazaki, T., 1681 Yamazumi, T., 1617 Yanagi, H., 159 Yanagi, N., 605, 735, 739, 743, 751, 1253, 1285, 1825 Yanagi, Y., 211 Yanagise, N., 531 Yanai, M., 263, 307 Yanaka, S., 783 Yang, J., 625 Yang, Z.-Q., 1475 Yano, Y., 855, 859 Yao, H., 217, 221,233, 497 Yasohama, K., 1763, 1779 Yasuda, M., 1431 Yasuda, T., 1643 Yasukawa, Y., 135, 795, 1009 Yasunaga, T., 1553 Yasuoka, H., 1537 Yatsuda, T., 1505 Yazaki, T., 243 Yazawa, T., 1029, 1113 Ye, J., 1635 Yin, Z.-Z., 639 Yokogawa, K., 1915, 1919, 1923 Yokoyama, S., 1133 Yoneda, E. S., 913, 917, 921 Yonenaga, Y., 747 Yoshida, K., 131, 135,811, 1731, 1771 Yoshida, N., 991, 1155, 1413, 1583, 1597, 1847, 1855, 1859 Yoshida, S., 457, 657 Yoshida, T., 739 Yoshikawa, K., 319, 481 Yoshikawa, M., 1095, 1293 Yoshimura, N., 319 Yoshinaga, S., 203 Yoshino, Y., 1907 Yoshitomi, J., 933, 991, 1003, 1867 Yoshizaki, R., 1385, 1435 Yoshizawa, S., 701 Yotsuya, T., 1201, 1591 Yu, J.-P., 247, 315, 407, 677 Yumura, H., 1855, 1859
A u t h o r Index Yuri, T., 1899 Yuyama, J., 267, 1083 Yuyama, M., 1735, 1783 Zadro, K., 1557 Zeng, D., 1475 Zeng, Z. J., 1037 Zhang, C., 1475
Zhang, C.-Q., 271 Zhang, L.-A., 343, 347, 613 Zhang, P., 555 Zhang, P.-X., 1447 Zhang, Z.-Y., 343 Zhao, L., 247, 275, 315 Zheng, J.-Y., 247, 275, 315, 677 Zheng, X. G., 1451, 1639
Zhou, G.-L., 229 Zhou, L., 1373, 1417, 1421 Zhou, L.-A., 1447 Zhou, S.-L., 625, 1635 Zhou, Y., 271 Zhu, S.-W., 291 Zhu, W., 1021
I9
