First U.K. National Conference on Heat Transfer. Volume 1.86

685 Pages·1984·27.224 MB·English

EFCE Event No. 308
First U.K. National Conference on
HEAT TRANSFER
Organised by the U.K. National Committee for Heat Transfer, which is a joint committee of
the Institutions of Chemical and Mechanical Engineers and includes a member nominated
by the Heat Transfer Society.
Held at the University of Leeds, 3-5 July 1984.
Organising Committee
Professor H C Simpson (Chairman) Strathclyde University
Dr G F Hewitt (Vice-Chairman) Harwell Laboratory
DrD Boland ICI Limited
DrTR Bott Birmingham University
Dr B N Furber NNC Limited
ProfessorWBHall Manchester University
DrPJ Heggs, Leeds University
Professor P N Rowe University College London
Mr E A D Saunders Whessoe Limited
Professor D B Spalding Imperial College London
THE INSTITUTION OF CHEMICAL ENGINEERS
SYMPOSIUM SERIES No. 86
ISBN 0 85295 170 1
l
PUBLISHED BY THE INSTITUTION OF CHEMICAL ENGINEERS
Copyright© 1984 The Institution of Chemical Engineers
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British Library Cataloguing in Publication Data
U.K. National Conference on Heat Transfer (1st: 1984 Leeds)
First U.K. National Conference on Heat Transfer.
— (EFCE event; no. 308) — (Institution of Chemical Engineers symposium
series, ISSN 0307-0492; no. 86)
1. Heat — Transmission
I. Title II. U.K. National Committee for Heat Transfer III. Series 536'.2 QC320
Pergamon Press ISBN 0-08-030281-5
Library of Congress No: 84-16695
11
PREFACE
Heat transfer is an area of of central concern to chemical engineers
and a topic in which they have a major interface with mechanical
engineers and others from a number of scientific disciplines.
The UK National Committee for Heat Transfer, together with the
Yorkshire Branch of the IChemE, has taken the initiative in deciding to
organise a series of national conferences to complement the existing
regular series of the International Heat Transfer Conference.
It is intended that the Leeds conference will be the first of a series of
UK National Conferences which will be held at four-yearly intervals
(1984, 1988, 1992 etc). Thus, for people working in the heat transfer
field (overseas as well as in the UK) there will be an opportunity to
present and discuss their work at a major conference every two years
(taking account of the International Conference which will be held in
1986, 1990, 1994 etc).
The papers in these two volumes have been subjected to peer review,
and have been modified to take account of reviewers' comments
where appropriate. The papers thus form a proper part of the archival
literature on heat transfer. The series of Proceedings arising from
these UK Conferences are intended to take their place as an
important component in the regular publication of reference material
on this continually developing subject.
in
CONTENTS-VOLUME 2 we
Session 11. ENHANCED HEAT TRANSFER
11.1 The prediction of heat-transfer performance in spirally fluted tubes: the 695
turbulent flow regime.
Barba, A., Gosman, A.D. and Launder, RE.
11.2 Heat transfer enhancement in shell/tube heat exchangers employing 707
electrostatic fields.
Poulter, R. and Miller, LA.
11.3 The Senftleben effect in the shell/tube heat exchanger. 717
Cooper, P. and Allen, P. KG.
11.4 Conjugate laminar and/or turbulent forced convection heat transfer from 725
rectangular fins.
Sunden, B.
11.5 The FLUFF code for calculating finned surface heat transfer . 7 35
Fry, C.J.
11.6 Air-water mist flow over extended surface tubebanks. 745
Simpson, H. C, Beggs, G. C and Raouf, M.I.N.
Session 12. TWO-PHASE FLOW AND BOILING
12.1 Prediction of flow pattern boundaries in horizontal two-phase flow. 761
Chisholm, D.
12.2 A simple model for estimating two-phase momentum flux. 773
Morris, S.D.
12.3 Saturation nucleate pool boiling - a simple correlation. 785
Cooper, M G.
12A Dryout on the shell side of tube bundles. 795
Schuller, R.B. and Cornwell, K.
12.5 Transient critical heat flux in flow boiling. 805
Whalley, P.B., Lyons, A.J. and Swinnerton, D.
12.6 Critical heat flux characteristics for vertical steam generating tubes with 817
circumferentially non-uniform heating.
Humphries, P., Dennett, G.F. and Scruton, B.
Session 13. NATURAL CONVECTION
13.1 Numerical solutions for developing combined convection between 829
uniformly heated vertical parallel plates.
Szpiro, O., Lewis, J.S. and Collins, M.W.
13.2 Heat transfer calculations including natural convection in porous insulation 839
and thermal radiation.
Oliver, A.J. and Stephenson, PL.
13.3 Measurements in a turbulent natural convection boundary layer. 849
Cheesewright, R. and Ierokipiotis, E.
13.4 Natural convection heat transfer measurements in a trapezoidal water- 1311
filled-plenum.
Birchenough, P.M., Bott, T.R. andLovegrove, PC.
13.5 Natural convection caused by passive conductors in stably stratified 857
fluids.
Quarini, G. L and Winters, K. H.
13.6 A study of natural convection in narrow annuli using small scale water 867
filled models.
So per, B.MH.
viii
Session 14. MEASUREMENT TECHNIQUES IN HEAT TRANSFER m*
14.1 Simulation of heat transfer in a partially blocked nine pin sub-channel 579
using the electrochemical technique.
Patrick, MA., Pembery, /. G.A. and Wragg, A.A.
14.2 Determination of heat transfer coefficients on film cooled flat surfaces 893
using the swollen polymer technique.
Saluja, CL, Lampard, D., Hay, N. and Burns, L
14.3 The application of the liquid crystal technique to the experimental 907
modelling of forced convection heat transfer in industrial heating processes.
Davies, R.M., Rhines, J.M. and Sidhu, B.S.
14.4 A technique for the measurement of local heat transfer coefficients using 919
copper foil
Watts, J. and Williams, F.
14.5 Insulation degradation errors in long, metal sheathed thermocouples: 933
calculations and measurements.
Burnett, R and Gentry, P.
Session 15. HEAT TRANSFER IN HIGH TEMPERATURE SYSTEMS
15.1 High temperature heat transfer: furnace performance assessment 951
Cooper, S. and Gibb, J.
15.2 Heat transfer from flames to convex surfaces. 969
Hemeson, A.D., Horsley, ME., Purvis, M.R. and Tariq, A.S.
15.3 Combined radiative and convective heat transfer in an enclosure. 979
Andrews, G.E., Gupta, ML and Mkpadi, M. C
15.4 Temperature field in solid forming an enclosure where heart transmission 989
by convection and radiation is taking place.
Bialecki, R., JNahlik, R. andNowak, AJ.
15.5 The radiation distribution factor: its calculation using Monte Carlo 1001
techniques and an example of its application.
Mohan, J.R. andEskin, LD.
Session 16. HEAT TRANSFER IN COMBUSTION SYSTEMS
16.1 Thermal modelling of fossil-fired boilers. 1013
Mobsby, J.A.
16.2 The effect of turbulence intensity on convective heat transfer from pre- 1025
mixed methane-air flames.
Hargrove, G.K. and Kilham, J.K.
16.3 Mathematical modelling of load-recuperative gas-fired furnaces. 1035
Tucker, RJ. and Lorton, R.
16.4 Transpiration cooling of gas turbine combustion chamber walls. 1047
Andrews, G.E. and Asere, A
16.5 Experiments and modelling of ignition on thermally thin plane-parallel 1057
pyrolizing solids under transient radiative heating.
Phuoc, T.X. and Durbetaki, P.
Session 17. CONVECTIVE HEAT TRANSFER
17.1 Low Reynolds number heat transfer in the entrance region of a parallel 1067
channel.
Ince, N.Z. and Hatton, A.P.
17.2 A numerical study of the influence of thermal stratification on forced 1077
convection heat transfer in sodium.
Hulme, G.
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page
17.3 Heat transfer for laminar flow in the entrance region of a smooth pipe. IQ91
Ingham, D.B.
17.4 The computation of momentum and heat transport in turbulent flow ^Q97
around pipe bends.
Iacovides, H. and Launder, B.E.
17.5 Full coverage impingement heat transfer: the influence of impingement ^15
jet size.
Andrews, G.E. and Hussain, CI.
Session 18. HEAT TRANSFER IN CROSS-FLOW
18.1 An investigation of the influence of freestream turbulence on the thermal 1125
characteristics around a single tube.
Hanarp, L. and Sunden, B.
18.2 The effect of longitudinal pitch upon the pressure drop across transverse 1137
rows of circular tubes.
Hollingsworth, M.A. and Mayhew, Y.R.
18.3 Convection heat transfer and flow resistance measurements in conventional 1147
and cross-inclined tube banks.
Hua, M, Le Feuvre, R.F. and Armstrong, C.
18.4 The prediction of shellside flow distribution and pressure drop in a 1163
shefl-and-tube heat exchanger.
Johnston, D. and Wills, MJ.N.
18.5 Pressure loss mechanism in resistances inclined to an air flow, with 1175
application to fintubes.
Mohandes, MA., Jones, T. V. and Russell, CM.R
Session 19. APPLIED HEAT TRANSFER
19.1 Boiling heat transfer at finned surfaces. 1191
Clayton, D. G.
19.2 'Appropriate* calculation methods for convective heat transfer from 1201
building surfaces.
Alamdari, F., Hammond, G.P. and Melo, C.
19.3 Heat transfer to protein concentrates. 1213
Obeng, E.D.A and Hoare, M.
19.4 Acoustic resonance in air cooled heat exchangers. 1223
Cow ell, T.A. and Davenport, CJ.
19.5 Hydraulic expansion of tube to tubesheet joints. 1235
Aldred, D.L
Session 20. INDUSTRIAL HEAT EXCHANGERS
20.1 Applications of novel, flexible, shell and tube heat exchangers. 1255
Roach, G.K and Wood, R.M.
20.2 Graphite heat exchangers in the process industries. 1263
Madner, RJ.
20.3 The plate heat exchanger: construction and design. I275
Kumar, H.
20.4 The flow distribution in plate heat exchangers. 2 291
Edwards, M.F., Ellis, D.L and Amooie-Foumeny, M
20.5 Plate heat exchangers for heat recovery. 1303
Cross, P.
x
STEADY STATE POST-DRYOUT HEAT TRANSFER IN A VERTICAL TUBE WITH LOW INLET
QUALITY
• * *
G. Costigan , A.W. Holmes , J.C. Ralph
A preliminary series of steady-state, post-dryout experiments
has been performed on a tubular Nimonic 90 test section
(8 mm bore, 9.5 mm O.D., 0.9 m long). Steady state was
achieved by using the "hot patch" technique. Measurements
were made in upflow and downflow at the following conditions:
pressure 2.5 - 3 bar inlet temperature 120°C - 125°C, mass
flux 50 kg/m2s to 200 kg/m2s. Valuable operating experience
and insight into the behaviour of hot patch type experiments
was obtained. The data indicate that for given inlet
conditions heat transfer in downflow is less efficient
than in upflow. For the range of conditions investigated,
it is thought that flow pattern changes may be affecting
the results.
INTRODUCTION
In recent years many experiments have been performed on flow boiling systems
in attempts to establish the mechanism of heat transfer immediately downstream
of a dryout location. The main emphasis of these experiments has been on the
low flows and inlet qualities considered typical of the reflood phase of a
postulated loss of coolant accident in a pressurised water reactor. These
experiments have demonstrated the complexity of the phenomena which occur in
the vicinity of a moving quench front.
A study of the sensitivity of a number of large computer codes, used for
the assessment of reactor safety, to variations in the methods of predicting
wall to fluid heat transfer concludes that:
"The amount of convective pre-cooling ahead of the quench front plays a
most important role in determining the rate of quenching." [1]
To predict the progression of a quench front accurately, knowledge of
heat transfer rates immediately prior to the arrival of the front is required.
Transient reflooding and quenching experiments result in a situation in which
the fluid quality on arrival at the quench front is constantly increasing
throughout the test. This is due to the transfer to the fluid of energy
stored and/or generated in the rewetted length of tube upstream of the front:
since this rewetted length is increasing as the quench front moves forward
the total energy added to the fluid between inlet and the front must be
integrated over this length. Whilst it may be argued that a transient experi
ment is more representative of a real reflood situation, the practical diffi
culties of accurately modelling the heat storage, generation and transport
properties of a nuclear fuel rod, and the changing fluid quality, make the
*
Engineering Sciences Division, AERE Harwell, Oxfordshire.
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Symposium Series No. 86
interpretation of the results more onerous. This problem is exacerbated by
the fact that there is strong evidence to suggest that a change in flow
pattern downstream of the dryout location takes place as the coolant becomes
saturated.
To avoid the problems associated with transient experiments the present
study used a development of the "hot patch" technique (originally proposed by
Groeneveld [2]) to produce steady state post dryout data for water in upflow
and downflow.
BACKGROUND
Experimental work
Experiments have been conducted on test sections as diverse as:
(i) liquid metal heated tubes and annuli [3]:
(ii) short, thick-walled metal test sections which are pre-heated and
then allowed to quench [4-6]:
(iii) long tubes which are also pre-heated and quenched [7, 8]:
(iv) tubes incorporating a large copper "hot patch" at inlet which
enabled near steady state post dryout conditions to be established
[9, 10].
These experiments have generally been performed at pressures between 1
and 4 bar. The data indicate that the post dryout heat transfer coefficient
falls with distance from the dryout location in the region immediately down
stream of the dryout point. Increases in heat transfer coefficient with inlet
flow rate and subcooling have also been reported, and a minimum in the heat
flux versus mass flux curve appears to exist at a low value of downflow velocity.
The available evidence [7, 8] suggests that the flow pattern which occurs
downstream of the dryout location is determined by the fluid quality immediate
ly prior to dryout. Thus if the quality just before dryout is negative
(corresponding to subcooled liquid) then inverted annular flow will occur:
however, if the quality at the quench front is positive then annular flow
below the quench front location will give way to dispersed droplet flow in
the post-dryout region (Figure 1). Kawaji [11] recently published the results
of a study of reflooding phenomena which include visual observations of flow
patterns and gamma densitometer measurements of void fraction at three axial
locations: these results support the above observations, and indicate that
the length of the inverted annular flow regime increases as the subcooling
at the quench front is increased.
The liquid core of the inverted annular flow pattern eventually breaks
down into a dispersed flow pattern as the relative velocity between the steam
and the liquid increases. Between the two extremes of inverted annular flow
and dispersed droplet flow Kawaji identifies a transition region, in which
large chunks of liquid from the liquid core are broken down into smaller
droplets. In this region a decrease in quality was found to produce an
increase in heat transfer. This is contrary to the predictions of dispersed
flow film boiling correlations, but supports the suggestions of Nelson [12].
In the positive quality pre-quench region Kawaji found that droplets
tended to be created by slugs of vapour periodically rising to the surface of
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the liquid and ejecting quantities of liquid into the vapour space. This is
in contrast to the wave entrainment observations of Ardron [13] and was
attributed to a shorter length of annular flow region upstream of the quench
front than was present in Ardron's experiment.
The quenching of hot surfaces by downflow of coolant was reviewed by
Butterworth [14] and the quench front flow pattern described was one in which
a falling film of liquid was thrown off the hot surface after a short length
of vigorous nucleate boiling. Experiments performed on the quenching of high
thermal inertia nickel blocks by water at low downflow velocities [6] revealed
that falling film type behaviour could exist in single tubes. If the water
downflow velocity was close to the rise velocity of a slug of steam in a
stagnant liquid, then a stationary slug of vapour was observed at the inlet
to the test section, around which the inlet water flowed. Condensation of
steam at the upstream end of the slug was presumably balanced by vapour rising
under buoyancy from the test section. At higher velocities and subcoolings
there was no evidence of the presence of any slug at inlet and it was thought
that inverted annular flow had been established in downflow.
Post Dryout Heat Transfer Prediction Methods
Yu [7] has correlated single tube reflooding data in terms of the velo
cities of liquid and of the quench front and the equilibrium quality rise due
to the passage of the front. The correlation accounts for differences between
subcooled and saturated coolant at the quench front, and predicts an exponen
tial decay of heat transfer coefficient with distance from the front.
Analytical models of post dryout heat transfer tend to concentrate on
either the inverted annular flow pattern or the dispersed flow pattern.
Inverted annular flow is most frequently treated in a similar way to the
original Bromley [15] analysis for film boiling on horizontal cylinders; one
such derivation - for film boiling on vertical surfaces - is given by
Andersen [16], The Bromley type analysis assumes that the liquid is saturated,
the vapour film is laminar, that heat transfer from the wall to the liquid
core occurs by conduction across this vapour film and that the vapour generated
increases the film thickness. In this way a local heat transfer coefficient
is derived as a function of the distance from the quench front 'z':
a = K{-K^ —PIv (PL1 " PJS£ Ah v* }lf" (1)
\ <Tw " V z
The value of K depends upon the assumed boundary condition at the vapour
liquid interface:
K = 0.5 for zero interface velocity
K = 0.707 for zero shear stress
It is customary to modify Ah to account for the superheat of the vapour thus:
Ah^ = Ah + A . c . AT (2)
v v s
The value of A depends upon the temperature profile in the vapour; for a
linear profile A = 0.5.
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Symposium Series No. 86
Chan [17] and Denham [18] have developed similar models of the laminar
film boiling inverted annular flow regime. Fung [19] has also developed a
model for inverted annular flow which covers both the laminar and turbulent
film regions. Each of the above models requires numerical solution. Neither,
as yet, has incorporated a method for predicting the heat transfer which occurs
as the liquid core breaks down into dispersed droplet flow. A further problem
with the prediction of heat transfer near the quench front arises when con
sidering how to partition the energy released at the quench front; this
affects the water subcooling and the vapour film thickness. Elias [20]
developed a solution for inverted annular flow which assumed that all the heat
released at the quench front formed vapour, and that thereafter heat was trans
ferred by conduction across a vapour layer of constant thickness to a sub-
cooled liquid core. Agreement with the data of Fung [9] was achieved by
adjusting the assumed film thickness.
A number of methods exist for the prediction of post-dryout heat transfer
with dispersed droplet flow. The majority of these have been reviewed by
Groeneveld [21], The empirical correlations are classified according to
whether they assume thermal equilibrium between the liquid and vapour (e.g.
[22]) or whether some degree of vapour superheating is permitted (e.g. [23]).
Non-equilibrium effects are recognised as being significant in this flow
regime and recent correlations attempt to take account of them [24-26].
Analytical or semi-empirical models (many of which are developments of the
basic model formulated by Bennett (27) are frequently used to predict post
dryout wall temperatures [28, 29].
EXPERIMENTAL WORK
Test Section
In an attempt to obtain reproducible steady state heat transfer data for
low flow, low quality water, the test section shown in Figure 2 was constructed.
A copper disc at each end of the test section was equipped with cartridge
heaters which were connected to an electronic temperature controller. The
controller could supply sufficient power to maintain a constant block tempe
rature. The high thermal diffusivity of the copper ensured that the power
was distributed as required over the short (2 cm) length of the copper block
adjacent to the bore of the tube. Thus the high critical heat flux, necessary
to initiate film boiling with subcooled water, was supplied locally at the
tube inlet. The remainder of the tube was heated by passing a high current
through it from a separate a.c. power supply (0-30 volts, 0-2500 amps). The
tube was made from Nimonic 90 and had the following dimensions: bore 8 mm,
O.D. 9.5 mm, length 900 mm. The bore of the copper blocks was protected from
oxidation by brazing into it an 8 mm bore, 8.5 mm O.D., stainless steel
cylinder which had been machined on each stub connecting pipe. Thermocouples
in the blocks allowed estimates to be made of the inlet and exit heat fluxes,
and the temperature of the tube wall was measured by thermocouples placed
at 2 cm axial intervals.
The test section was mounted in the Harwell low pressure steam/water
loop described in [6]. The instrumentation was connected to a data logger
which was in turn connected to a desk top computer, and the data was stored
on floppy disks for subsequent analysis.
Procedure
The desired pressure, flow rate and inlet water temperature were set by
circulating water through a by-pass before beginning a test run. The copper
4