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Flow Boiling Heat Trader In The Quenching Of
A Hot Surface Under Reduced Gravity Conditions
Jason Jianxin Xu
A thesis submitted in conformity with the requirements for the Degree of Doctor of Philosophy
Graduate Department of Chernieal Engineering and Appiied Chernistry in the University of Toronto
@ Copyright by Jason Jianxin Xu 1998
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For My Parents and My Wife
Abstract
Flow Boihg Heat Tramfer In The Quenching Of
A Hot Surface Under Reduced Gravity Conditions
By: Jason Jianxin Xu
A thesis submitted in conformity with the requirements
for the Degree of Doctor of Philosophy
Graduate Department of Chemical Engineering and Applied Chexnistry
in the University of Toronto
An experimental set-up, which combined a new state of the art micro-sensor for
instantaneous measurements of heat flux and surface temperature, was designed, constmcted
and used to study the effects of gravity, as well as inlet liquid flow rate and subcooling on
rewetting of a hot horizontal surface. The experiments were conducted by injecting liquid
RI13 and PF5060 into an initially dry, heated channel, which was 40 mm wide, 5 mm high
and 200 mm long, on the ground and in reduced gravity aboard the parabolic aircraft, KC- 135
and DC-9 of the NASA. The measurements showed large instantaneous fluctuations in heat
flux and surface temperature following the onset of rewetting, even after the maximum heat
flux was passed, where the heat transfer mode changed from transition boiling to nucleate
boiling. Heat flux and surface temperature data showed synchronized responses indicating
sufficiently fast response of the sensors and the reliability of the measurements. The boiling
curves covering film, transition and nucleate boiling regimes were obtained during quenching
and analyzed. The heat transfer characteristics in each boiling mode, as well as rewetting
temperature, quench velocity, liquid-solid contact frequency in transition boiling and
maximum heat flux were examined in detail for different gravity Ievels, inlet liquid flow rate
and subcooling. The quench velocity and rewetting temperature were found to decrease for
RI13 but only showed very slight decreases for PF5060 in reduced gravity. A peak in the
liquid-solid contact frequency curve was found at wall superheats of 107 - 1 18 OC for R113
and 65 - 83 O C for PF5060 in both gravity conditions. The maximum heat flux for both fluids
decreased in reduced gravity except for R 1 13 at high flow rate.
Acknowledgments
1 would like to present my sincerely thankful feelings to dl the people Who had given
me help andor assistance dunng the progress of the study:
Dr. M. Kawaji, far his invaluable support, guidance and encouragement, and for many
valuable suggestions, especially those in improving the design of the experiments and
analysis of data during the course of this snidy;
Dr. D.C.S. Kuhn and Dr. S. Chandra, for their valuable suggestions during the reading
cornmittee meetings, which benefit my study very much.
The University of Toronto, Mitsubishi Electric Co. and Canadian Space Agency for the
financial support. The latter also provided the flight opportunities aboard the KC-135 and
DC-9.
Dr. K. Adham-Khodaparast, for his cooperation in the design and construction of the
experimental apparatus, and for his help in carrying out the experiments;
Mr. R. Lui, for design advice and constructive criticism, and for his help during the KC-135
campaign in Houston;
Canadian astronauts, Mr. M. Garneau and C. Hadfield, for their assistance during reduced
gravity experiments aboard the KC- 1 35 aircraft in Houston;
Mr. L. Vezina and S. Desjardins of Canadian Space Agency, for their organizing the flights
aboard the KC- 135 and the DC-9, respectively.
Mr. L. Rogers and the rest of the machine shop staff, for design advice and cnticism,
exceptionai professional workmanship and their extraordinary assistance during the shipment
of the experirnentai apparatus;
Mr. D. Tomchyshyn of Electrical Services, for his invaluable suggestions and assistance;
Mr. F. Shue, for his advice on superior construction and design of the test section;
The students in the lab, for their patience, understanding and assistance in various aspects
of the project;
My wife, Qingping Liu, has been the strongest supporter of my study. Therefore, it is a
pleasure to dedicate the thesis to her and Our beloved parents.
Table of Contents
Abstract
AcknowIedgments
Table of Contents
List of Tables
List of Figures
Nomenclature
1.0. Introduction and Literature Review
1.1. Background
1.2. Literature Review
1.2.1. Rewetting under Normal Gravity
1.2.2. Rewetting and Related Studies under Reduced Gravity
1.3. Research Objectives
2.0. Experimentai Apparatus
2.1. Test b o p and Components
2.2. Heat Flux Micro Sensor
2.3. Test Section
2.4. Condenser
2.5. Data Acquisition System
2.6. Visualization of Quenching Experiments
2.6.1. Quenching of a hot surface with subcooled R 1 13
2.6.2. Quenching of a hot rectangular quartz tube with subcooled water
3.0. Reduced Gravity Experiments and Procedure
3.1. KC- 135 and DC-9 Aircraft
33. Operational Procedure
3.3. Condenser Performance and Flow Rates
3.4. Quenching Experiments
v
vii . *.
V l l l
xiii
1
1
2
2
6
13
14
14
16
20
25
28
29
29
3.5. Heat Flux and RTS Sensors Performance
4.0. Quenching Experimental Resuits
4.1. Transient Heat Fiux and Surface Temperature Characteristics
4.2. Boiling Curves during Quenching
4.3. Fiow Boiling Visualization
4.3.1. Quenching experiments with heat flux indicator
4.3.2. Experiments on quenching of a quartz tube with water
5.0. Film Boiling, Rewetüng Temperature and Quench Velocity
5.1. Film Boiling
5.2. Rewetting Temperature
5.3. Quench Velocity
6.0. Transition Boiling Heat Transfer
6.1. The Mechanism of Transition Boiling
6.1.1. Background
6.13. Past modeling efforts under pool boiling conditions
6.13. Past experimental work on liquid-solid contact
6.1.4. Liquid-solid contact frequency results for R 1 13
6.1.5. Liquid-solid contact frequency results for PF5060
6.2. Transition Boiling Heat Transfer of R 1 13
7.0. Maximum Heat Flux and Nucleate Boiling Heat Transfer
7.1. Maximum Heat Flux
7.2. Nucleate Boiling heat transfer of RI 13
8.0. Cornparison of R113 and PF5060 Results
9.0. Conclusions and Recomrnendations
References
Appendix I(1) The Effect of the RFMS Disk Material
Appendix 1(2) The Effect of g-jitter on Quenching and Boiling Heat Transfer Characteristics
Appendix I(3) The Effect of Surface Tension
Appendix II Uncertainty Analysis
List of Tables
Table No. Description Page
The List of p-g and 1-g Runs (KC- 135 carnpaign).
The List of y g and 1-g Runs (DC-9 carnpaign).
Film boiling heat transfer coefficients in p-g and 1 -g experiments (R 1 13).
Quench velocity for RI 13 in p-g and 1-g experirnents. 85
The factors in correlation (6.5) for the mns in p-g and 1 -g experiments (R 1 13).
The constants in equation (7.12) for p-g and 1 -g experiments (R 1 13).
The data of b obtained by other researchers. 121
Fluid and thermophysical properties of R 1 13 and PF5060 122
The experimental conditions for R 1 13 on stainiess steel disk. 1
Cornparison of rewetting temperature between two HFMS disks 1
Cornparison of quench velocity on two HFMS disks n Cornparison of the maximum heat flux on two HFMS disks III
vii
List of Figures
Figure No.
Typical quenching curve during rewening of a hot surface.
Schematic of the test loop.
Schematic of the heat flux micro sensor.
The working principle of the heat flux micro sensor.
Schematic of the test section.
Schematic of the alurninum case cover.
A whole view of the test section.
Schematic of the condenser.
The 80w chart of the data acquisition system.
The heat flux indication system consisted of heat flux micro-sensor, amplifier, LCD and fiber glass.
The schematic of the test loop in quenching of a quartz tube with water.
Typical accelerations during the KC- 1 35 flights.
The reservoir full of liquid in normal gravity and hyper-gavity.
The blocked bubble columns at low flow rates in pg.
The blocked bubble columns at low flow rates in p-g.
The blocked bubble columns at intermediate flow rates in p-g.
The blocked bubble columns at high flow rates in p-g.
The collapsed bubbles entering the flow loop in p-g.
Typical flow rate profile for one run in p.-g (KC- 135).
Typical flow rate profile for one run in p.-g (DC-9).
The calibration curve for H F M S provided by the manufacture.
The calibration curve for the RTS with copper disk.
The calibration curve for the RTS with stainless steel disk.
A typical transient heat flux history during rewetting of RI 13.
Figure No.
4.2.
4.3
4.4.
4.5.
4.6.
4.7.
4.8.
4.9.
4.10.
4.1 1.
4.12.
4.13.
4.14.
4.15.
4.16.
4.17.
4.18.
4.19.
5.1.
5.2 (a).
5.2 (b).
Description
A typical transient surface superheat history during rewetting of R113.
Typicai power spectrum density of heat flux before R 1 13 injection.
Typicai power spectrum density of wail superheat before R113 injection.
Synchronized response showed by heat flux and surface temperature.
A typical transient heat flux history during rewetting of PF5060.
A typical transient surface superheat history during rewetting of PF5060.
Synchronized response showed by heat flux and surface temperature.
The effect of time average on boiling curve for R 1 13.
Boiling curves measured in 1-g for R 1 13.
Comparison of boiling curves measured in p-g and 1 -g for R 1 13.
Boiling curves in 1-g with low and high inlet subcooling for PF5060.
Boiling curves in p-g and 1 -g with high inlet subcooling for PF5060.
The discussion of boiling mechanism.
The quenching of RI13 image at film boiling regime.
The quenching of RI13 image at transition boiling regime.
The quenching of R113 image at maximum heat flux.
The quenching of R113 image at nucleate boiling regime.
The vapor film broke behind the quench front.
Typical film boiling heat transfer coefficient profile.
Comparison of film boiling heat transfer data for RI13 with the results predicted by equations (5.1) to (5.4), Uin = 0.38 &S.
Cornparison of film boiling heat transfer data for R 1 13 with the results predicted by equations (5.1) to (5.4), U, = 0.88 mis.
Page
49
50
50
53
54
55
55
57
58
58
Figure No.
5.3.
Page
The effect of initiai wdl superheat on rewetting temperature for R 1 1 3.
The rewening superheat for RI13 in p-g and 1-g.
The rewetting superheat for PF5060 in p-g and 1-g.
Cornparison of rewetting temperature for RI13 and PF5060.
The parameter, ATm/ATw VS. G for R 1 1 3 and PF5060.
Determination of quench velocity.
The quench velocity for R 1 13 in p-g and 1 -g.
Quench velocity for PF5060 in p-g and 1 -g.
Comparison of the quench velocity of R 1 13 and PF5060.
The ratio of quench velocity to inlet velocity for R 1 13 and PFS060.
Typical magnitude of heat flux fluctuations for R 1 13.
Typical magnitude of temperature fluctuations for R 1 13.
Typical power spectra of q and Tw fluctuations for R 1 13.
The effects of wall superheat and mass flux on liquid-solid contact frequency for R 1 13.
Heat flux fluctuations for PF5060 in transition boiling regime in p-g.
Heat flux fluctuations for PF5060 in transition boiling regirne in 1 -g.
Typical magnitude of heat flux fluctuations for PF5060.
Typical power spectra of heat flux fluctuations for PF5060.
Liquid-solid contact frequency for PF5060 in p-g.
Liquid-solid contact frequency for PF5060 in 1-g with high subcooling inlet.
Liquid-solid contact frequency for PF5060 in 1-g with low subcooling inlet.
Comparison of liquid-solid contact frequency for PF5060 in p-g and 1 -g with low inlet flow rate.
Comparison of liquid-solid contact frequency for PF5060 in p-g and 1-g with high inlet flow rate.
Definitions of boiling heat transfer regions.
Transition boiling heat transfer for R 1 13 measured in 1-g.
Comparison of transition boiling heat tmnsfer for RI13 in p-g and 1-g.
The maximum heat flux data for R 1 13 in both gravity conditions.
The maximum heat flux data for PF5060 in p-g and 1-g.
The qJGhlw for R 1 13 varies with We in both gravity conditions.
The q d G h l w for PF5060 varies with We in both gravity conditions.
The effect of flow on nucleate boiling in 1-g with subcooled inlet.
7.5 (a).
The effect of flow on nucleate boiling in 1-g with saturated inlet.
7.5 (b).
The effect of flow on nucleate boiling in p-g with subcooled inlet.
7.5 (c).
The effect of subcooling on nucleate boiling in 1-g. 7.6 (a).
7.6 (b). The effect of subcooling on nucleate boiling in p-g.
The effect of gravity on nucleate boiling heat transfer.
Comparison of nucleate boiling in the absence of subcooling and gravity.
Comparison of the boiling curves of R 1 13 and PF5060.
Comparison of maximum heat fluxes with subcooled inlet in 1-g.
Comparison of maximum heat fluxes with subcooled inlet in p-g.
Cornparison of subcooled q&Ghtv vs. We with the prediction in 1 -g.
Comparison of subcooled qJGhiv vs. We with the prediction in p.-g.
The quench velocity for a copper disk HFMS.
Figure No. Description
A.2. The typical power spectrum of gravity level Fiuctuations aboard the KC- 135.
A.3. The typicd power spectrurn of gravity level FIuctuations aboard the DC-9.
Page
v
v
xii
Nomenclature
hl"
n
P
PSD
normal gravity conditions
reduced gravity conditions
acceleration, m/s2, or constant
constant, or area m'
constant
constant
constant
heat capacity, kJkgK
diameter, m
liquid-solid contact frequency, Hz
fraction of wetted area
local liquid contact-time fraction
gravitational acceleration, 9.8 mis'
mass flux, kg/m2s
boiling heat transfer coefficient, kw/rn2 K
latent heat of vaponzation, k l k g
superficiai velocity, d s
Jakob number, dT.,&pJhfV
thermal conductivity, WlmK
charactenstic length or the length of plate, m
constant
constant
pressure, MPa
power spectral density
heat flux, kw/rn2
temperature, OC
velocity, mm/s
heat flux input voltage, Volt
... Xll l
Weber number, L G ~ / ~ O
quaiity
elevation, m
Greek Letters
A change in quantity
6 wall thickness
Âo criticai wave length, m
P dynamic viscosity, kg/ms
P density, kg/m3
O surface tension, N/m
Subscripts
ar
Ber
Bereson
Bromly
C
CHF
fb
h
in
1
m
max
nb
O
9
RTS
apparent rewetting
Bereson's correlation
Bereson's correlation
Bromly's correlation
critical conditions
critical heat flux
film boiling
h ydraul ic
inlet conditions
liquid properties
minimum film boiling conditions
maximum heat flux conditions
minimum film boiling conditions
nucleate boiling
initial wall conditions
quench
resistance temperature sensor rneasurement
xiv
sat
sub
Overbar
rewetting
thermocouple measurement
transition boiling
saturation conditions
subcooled conditions
vapor properties
wall properties, wall conditions
x dimension
y dimension
z dimension
dimensionless fraction
CHAPTER 1
Introduction and Literature Review
1.1. Background
Surface quenching or rewetting refers to the establishment of a continuous liquid
contact with a dry and hot solid surface and rapid cooling of the surface due to boiling heat
transfer. There are many important manufacturing processes that can occur under reduced and
normal gravity conditions involving quenching or rewetting of hot surfaces by a liquid. The
quenching process, such as in material processing, is cornrnonly encountered due to high
initial temperatures of surfaces compared to the boiling point of the fluid. In recent years, the
interest in quenching process has increased rnainly in connection with the safety analysis of
nuclear power reactors during hypotheticd loss of coolant accidents. More efficient heat
transmission in thermal management systems aboard the International Space Station, safe
refueling of space transfer vehicles (STV) with liquid hydrogen or oxygen propellants, and
transfer and storage of cryogenic fluids on the ground and in space, al1 require a knowledge
of quenching heat aansfer characteristics which could be significantly different under
reduced and normal gravity conditions.
There are many publications on quenching of hot surfaces under normal gravity, most
of which have focused on rewetting and reflooding of hot and dry vertical tubes. The
quenching data for horizontal channels are scarce even under normal gravity, and there have
been few quenching experiments perfonned for either horizontal or vertical tubes under
reduced gravity conditions. In the present study, a test loop and a test section incorporating a
new state-of-the art micro heat flux and surface temperature sensors were constructed and
transient quenching experiments were conducted at different gravity levels in order to
investigate the effects of gravity, liquid flow rate and subcooling on the rewetting
phenornena. In the next sections, the relevant literature is reviewed and the objectives of the
present work are described.
1.2. Literature Review
12.1. Rewetting under Normal Gravity
With the increasing requirements for removal of higher heat fluxes from heated or hot
surfaces, the snidy of boiling heat transfer has attracted more attention in the field of heat
transfer in recent years. In particular, rewetting of a hot surface is important in various
cooling processes involved in loss-of-coolant accidents of nuclear reacton. steel making and
other industrial processes.
On rewetting of hot surfaces, the previous studies have been Iargely concemed with
the collapse of a vapor film during pool film boiling in order to undentand the rewetting
mechanisms. Much experimental and theoretical work has been performed on the rewening
of heated vertical tubes in connection with the postulated loss-of-coolant accidents in
Pressurized Water Reactors ( P m ) .
Dunng rewetting, a hot surface experiences three boiling regimes: film boiling,
transition boiling and nucleate boiling, as illustrated in the boiling curve (Figure 1.1), which
is a plot of surface heat flux, q, against wall superheat (Le., difference between the wall
temperature and saturation temperature of liquid or Tw-Tm). Two transition points on the
boiling curve are the Minimum Heat Flux or quenching point and the Maximum Heat Flux or
Critical Heat Rux ( C m , the values of which are important in the studies of boiling heat
transfer and rewetting of hot surfaces. When a liquid droplet is placed on a hot surface, vapor
generated at the vapor-liquid interface forms a thin vapor film preventing the liquid from
contacting the hot surface. This boiling mode is termed film boiling and heat transfer is
dominated by conduction, or convection of heat from the hot wal1 to vapor and radiation of
heat through the vapor film. As the surface is cooled to the quenching point where heat flux
drops to a local minimum value, the vapor film becomes unstable and collapses. Then
quenching starts with large heat flux fluctuations due to intermittent liquid-surface contacts
and the process enters the transition boiling regime, in which the surface is locally dried out
and rewetted repeatedly until stable nucleate boiling is established.
Pool film boiling and flow film boiling heat transfer processes under steady state
conditions have been extensively studied in the past several decades and many theoretical
models showing good agreement with experimental data have been proposed. A
comprehensive review of film boiling heat transfer was presented by Carey (1 992).
O 5 10 15 20
Time (second)
160 -.
120 -
80 -
40 -
Figure 1.1. Typical quenching curve during rewetting of a hot surface.
I : : I I
:- Rewetting Run MG-D23
As the quench front progresses dong a flow channel, different two-phase flow
patterns appear depending on the flow rate, initial wall temperature and Iiquid subcooling.
For example, an annular flow pattern was observed for Iow injection rates of saturated liquids
and inverted annular flow pattem for higher injection rates of subcooled liquids in a vertical
tube by Kawaji et al. (1983) and others.
Numerous theoretical and experimental studies on the rewetting of vertical and
horizontal hot surfaces have been reviewed by Abdul-Razzak (1990) and Westbye (1992).
Three different controlling mechanisms have been suggested for forced convection rewetting:
(1) coilapse of a vapor film, (2) axial conductioncontrolled rewetting, (3) dispersed droplet
flow rewetting. The effects of many parameters such as inlet liquid subcooling, flow rate and
initial wall temperature, have been reported in many studies as being similar in both vertical
and horizontal systems. These studies indicate that the rewetting velocity increases with
decreasing initial wall temperature and channel wall thickness, and increasing liquid
subcooling, flow rate, systern pressure and surface roughness. Heat transfer increases in al1
boiling modes with increasing subcooling and flow rate.
Chan and Banerjee (1981) conducted experiments on refilling and rewetting of a hot
horizontal Zircaloy-2 tube with 19.6mm O.D. and 0.898 mm wall thickness using water at
atmospheric pressure. They found that gravitational effects were important and could lead to
flow stratification. The rewetting front was preceded by a liquid layer that was supported by
film boiling and termed a "liquid tongue". Significant precunory cooling was observed at the
tube's bottom surface due to the presence of this tongue. The effects of initial wall
temperature and inlet flow rate on quench velocity were consistent with the results obtained
by others in vertical rewetting experiments.
Recently, Huang et al. ( 1994) reported three different types of quenching for a vertical
hollow copper tube of 50 mm length, 10 mm LD. and 32 mm O.D. The experiments were
conducted at different pressures (P) ranging from 0.1 to 1.0 MPa, mass flux (G) from 25 to
500 kg/m2s and inlet subcooling (ATrub) from 5 to 50 K to study the transient effects in a
quenching process. Thirty-two sheathed NiCr-Ni micro-thermocouples, monitored at a
sarnpling frequency of 10 Hz, were press-fitted inside the tube to determine the transient
temperature field. They found that the transient effects were not obvious for inlet conditions
of G < 300 kg/m2s, P < 0.7 MPa and ATsub < 15 K. Beyond this range, however, the
quenching curves lay below the steady-state boiling curves in transition boiling regime, with
the minimum film boiling temperature king correspondingly lower. Also they found that
there existed a significant difference between two types of quenching, quenching of a dry, hot
tube and quenching a tube with the power cut off after establishing a stable inverted annular
flow.
Barnea and Elias (1994) experimentally and theoretically studied flow and heat
transfer regimes during quenching of a heated, vertical circula. channel with 44.8 mm I.D. at
pressures ranging from 0.1 to 0.5 bar, and for the initial wall temperature between 350 and
600 O C , and two liquid inlet temperatures: 30 and 60 OC. Water was used as a coolant, and
tube wall temperature, liquid and vapor temperatures in the flow channel were measured by
ungrounded type K thermocouples. The results indicated that the ratio between the quench
velocity and the inlet liquid velocity varied between 0.2 and 0.8. At the quench front a small
fraction of liquid formed a thin layer in contact with the wall and the average thickness of the
liquid layer could be detemiined from the rate of heat transfer from the wall. They found that
there existed four heat transfer zones dong the wall, defined as the high surface temperature
zone, forced convective subcooled boiling zone, "wet" and "dry" transition zones and
subcooled inverted annular flow zone. In the inverted annular Row zone the vapor film
thickness was almost constant in the fint 30 to 50 mm downstream of the quench front. From
the measurement of void fraction, the vapor film extending downstream of the quench front
was observed and it revealed no cornplete collapse of the vapor film at the quench front.
In order to compare boiling heat transfer results from quenching experiments with
those from steady state boiling experiments, Bergles and Thompson (1970) quenched metal
spheres in quiescent water and other Buids. Their results indicated that the transient boiling
curve for water was considerably shifted to the higher wall temperatures, particularly in the
nucleate boiling regime. Similar results were obtained by Jacob and Dougall (1978) in flow
quenching of a thin-walled heater tube. Peyayopanakul and Westwater (1978) performed a
preliminary study on the limitations of the transient quenching method for pool boiling on a
copper cylinder, and Westwater et al. (1986) did a similar study on pool boiling around a
sphere and over a horizontal flat plate facing upward. They found that the dimensions of the
test section mainly affected the boiling curve, in other words, the establishment of the quasi-
steady-state depended upon the thermal capacity of the test section. Also the properties of the
metal were found to affect the boiling curve.
Moreover, houe and Tanaka (1991) performed some experiments on rewetting of a
vertical, 10 mm I.D. tube by injecting R-113. They rneasured steady-state boiling curves and
transient quenching curves using a copper block for mass flux ranging frorn 412 to 1466
kglm's, inlet flow quality from 0.6 to -0.29 and pressure from 0.26 to 0.30 MPa The steady-
state boiling curves obtained showed good agreement with quenching curves due to a high
heat capacity of the test section and both were greatly affected by the inlet flow qudity (mas
flow rate of vapor divided by the total mass flow rate). The rewetting mechanism, which was
only slightly affected by m a s flux, changed with the flow quality; rewetting due to dispersed
80w was observed for quality higher than 0.2, and rewetting due to inverted annular flow for
quality less than 0.1. Rewetting temperatures were also measured in dispersed flow and
inverted annular flow film boiling, and were constant in the saturated flow region but
increased with liquid subcooling.
1.2.2. Rewetting and Related Studies under Reduced Gravity
Although the studies on two-phase tlow and boiling heat transfer under reduced
gravity have received considerable attention in the past two decades with the increasing
applications in space, the literahue is still scarce. The behavior of two-phase flow systems
under reduced gravity have been reviewed a decade ago by Rezkallah (1988). It was
concluded that in general, segregated and intermittent flow regimes are restricted to a more
narrow range of flow qualities and liquid velocities, while a dispened flow regime covers
wider ranges of quality and velocity. Also under reduced gravity, different flow regimes such
as "frothy annular" can occur, and the two-phase pressure &op generally increases. Recently,
Karnp et al. (1995) studied the radial distributions of void fraction, velocity and turbulence
intensity in upward (lg), downward (-lg) and reduced gravity aboard the Caravelle aircraft in
bubbly flow in a pipe. They found that void coring also existed in reduced gravity but the
void fraction profile was flatter than that for downward flow in normal gravity. Jayawardena
et al. (1997) proposed a general flow pattern map based on dimensionless parameters for
rnicrogravity two-phase flows, which is different from the previous flow maps using gas and
liquid superficial velocities.
Rite and Rezkallah (1994, 1997) studied convective heat transfer in a vertically,
upward co-current flow of water and air through a circular tube aboard the NASA's KC- 135
parabolic aircraft. They found that the heat transfer coefficients in reduced gravity were lower
than in normal gravity at low liquid and gas velocities and this trend was reversed at higher
liquid or gas velocities. Also, reduced gravity had a tendency to lower the heat transfer
coefficient by as much as 50% at the lowest flow qualities in the bubbly and slug flow
regimes. As the flow quality increased and the flow regime changed to annular flow, the
differences in heat transfer coefficient between the normal gravity and reduced gravity were
much less significant.
Recently, Issacci et al. (1995) conducted a literature survey on two-phase flow and
heat transfer involving nucleate pool boiling or forced convective boiling under reduced
gravity conditions. They concluded that the flow regime maps developed for two-phase flow
under low gravity conditions were still not fully established at present. Surprisingly, they
found that most low gravity maximum heat flux data are still those summarized by Siegel
(1967). Also, the siudies on 80w boiling under low gravity are very few and none of them
have investigated the critical heat flux.
The boiling experiments under microgravity conditions started with small drop tower
tests. Merte and Clark (1964) performed pool boiling experiments by immersing a 2.5 cm
diameter copper sphere in liquid nitrogen as a transient calonmeter using a 10 meter high
&op tower with a drop time of 1.4 seconds. Fractional gravity down to 0.0 1 g was obtained
using appropriate counter-weights and the expenrnental results showed that the film boiling
heat transfer coefficient was proportional to gl", the critical and minimum heat flux followed
a g1f4 dependence, and that the nucleate and transition boiling regimes were unafTected by
changes in gravitational acceleration.
However, Siegel and Keshock (1965) studied the effect of gravity on pool boiling
using a 3.8 rn drop tower and a platinum wire of 0.5 mm diameter in horizontal and vertical
orientations. They found that the heat transfer coefficient increased for two of their three test
fluids for the horizontal wire and decreased for the vertical orientation of the wire in
rnicrogravity. Siegel (1967) also reviewed the early studies on microgravity pool boiling.
Straub et al. (1990) summarized their 15 years of rnicrogravity pool boiling
experiments as showing a small effect of gravity on nucleate pool boiling, which was
contradictory to many pool boiling models which strongly rely on the effect of buoyancy on
nucleation and vapor bubble release. They also found that cntical heat fluxes were higher
than the values calculated by the correlations of Zuber (1959) and Lienhard and Dhir (1973),
and film boiling heat transfer decreased with gravity, but remained nearly constant for lower
g levels.
Recently, Ervin et al. (1992) and Ervin and Merte (1993) conducted transient pool
boiling experiments using RI13 in a drop tower to study nucleation and pool boiling
mechanisms in microgravity. They used a transparent gold film sputtered on a quartz
substrate, which worked both as a heating element and resistance temperature sensor at the
sarne time and also pennitted viewing of the boiling process from beneath the heated surface.
A new energetic type of boiling spread was observed in microgravity due to an interfacial
instability driven by a large mass flux across the wnnkled liquid-vapor interface.
Shortly later, Merte ( 1994) reported the results of pool boiling experiments conducted
aboard the Space Shuttle with the sarne surface as was used in the drop tower tests described
above. Subcooled boiling during long periods of microgravity was found to be unstable. The
heating surface was found to dry out and rewet, and the average heat transfer coefficients
during the dryout and rewetting periods were found to be about the sarne. Very recently,
Merte and Lee ( 1997) summarized their pool boiling experimental results conducted in three
Space Shuttle missions, STS 47, 57 and 60, using the same experimental apparatus. They
found that the absence of buoyancy resulted in the onset of boiling at low levels of heat flux
not othenvise possible in earth gravity. With the highest heat flux imposed on the heater
surface during the missions, heterogeneous nucleation was observed through the transparent
heater surface. However, a special form of homogeneous nucleation, which was termed as
quasi-homogeneous nucleation, occurred at various random locations followed by violent
vapor bubble growth rates when lower heat €luxes were imposed.
In addition to the study of pool boiling in reduced gravity in the United States and
Europe, two groups from Japan have studied the mechanisms of pool boiling using single and
binary mixture fluids in reduced gravity. Abe et al. (1994) conducted pool boiling
experiments with non-azeotropic water-ethanol mixtures in a 10-second drop tower.
Nucleation of bubbles on a transparent Pyrex plate with a thinly coated transparent heater
film was observed from side and below at a certain angle. The observations showed that the
hlly grown bubbles imrnediately detached but stayed at a small distance away from the
heater surface. The liquid supply to the base of each as-yet undetached bubble was found to
be intensive, possibly because of a Marangoni fiow effect. Also, large coalesced bubbles
formed above an array of small primary bubbles attached to the heater surface and continually
absorbed the small bubbles at higher heat inputs. The heat transfer measurements showed
enhancement in nucleate boiling but reduction in CHF in reduced gravity. Using sirnilar
experimental setup, they performed a series of pool boiling experiments using subcooled n-
pentane, CFC-113 and water aboard the Caravelle 234 aircraft (Oka et al., 1995). They found
significant reduction of in CHF but very slight changes in nucleate boiling heat transfer at
low heat fluxes for n-pentane and CFC-113. Opposite results were obtained for water and this
difference was believed to be attributed to a considerable difference in surface tension and
wettability between the organic fluids and water.
Ohta et al. (1997, 1998) reported pool boiting results obtained from the NASDA TR-
1A sounding rocket experiment. The boiling process with the systern pressure varying from
0.0 1 to 0.48 MPa on a sapphire glass disk using ethanol was viewed from the side and below
the heater surface. Different from many sirnilar heating and measuring techniques used
previously, they coated the back surface of the sapphire disk with a transparent heating film
so that the platinum temperature sensors, which were directly deposited on the boiling
surface, were electrically insulated from the heater. Steady state nucleate boiling was
achieved except in the case of low liquid subcooling during a six minutes period of high
quality reduced gravity. They observed that the behavior of small primary bubbles, existing at
the base of the large coalesced bubble, dorninated nucleate boiling for high heat inputs or at
low liquid subcooling. They also found that there existed two opposite trends of heat transfer
enhancement and deterioration depending on the behavior of the evaporating microlayer and
that of the dry patch at the base of the large bubble. Using the measured liquid film thickness
data, they developed a new mode1 to predict the surface heat flux.
Most of the microgravity experiments on boiling to date have been concemed with
pool boiling situations in order to clariw the basic mechanisms of nucleate boiling, bubble
generation and boiling incipience. On the other hand, flow boiling studies under rnicrogravity
are much more useful for practical applications, however, there are few studies reported in
the past.
One of the earliest studies on flow boiling under reduced gravity conditions is that of
Cochran (1970). The experiments were conducted in drop towers with water flowing over a
thin Chrome1 flat plate heater to study boiling process near inception. In cornparison to
normal gravity tests, it was found that bubbles tended to stay on the heating surface so that
they could become large enough to coalesce with the neighboring bubbles and acquire
irregular shapes in microgravity .
Recently, with the help of parabolic flight aircraft which can produce reduced gravity
periods longer than those in drop towers, Lui, Kawaji and Ogushi (1994) conducted a series
of experiments on flow boiling of RI 13 in a thin-walled stainless steel tube in horizontal
orientation aboard the KC-135 of NASA. Their results showed that the entire heat transfer
coefficient increased for subcooled, low quality, flow (nucleate) boiling during reduced
gravity by a factor of 5 to 20% over the normal gravity conditions. They attributed the
enhancement to the changes in phase distribution, and the greater movement of the vapor
bubbles generated on the heated tube surface, which could induce more localized turbulence
near the heater surface.
Saito et al. (1994) also studied flow boiling of water on a heater rod placed in a square
channel. The experiments were perforrned in the Japanese low gravity aircraft (MU-300).
They dso found that the nucleate boiling heat transfer coefficients slightly increased in the
flow direction but the magnitudes of the heat transfer coefficient were about the sarne as
those in normal gravity.
Recently, Wang et al. (1996) compared subcooled pool and forced-convection
nucleate boiling of R113 on a semitransparent gold-film heater in normal gravity and reduced
gravity, which was created by drop tower. Their measurements and observations showed that
higher heat transfer coefficients were achieved under reduced gravity conditions for both
types of boiling modes except for the case with the highest heat flux input (7.58 ~ lcm' ) , in
which the heat transfer coefficients declined sharply due to the formation of vapor slugs
above the heater surface and drying-out of some portions of the heater surface. The effect of
flow on boiling heat transfer enhancernent was only limited to the lowest heat flux input
(2.88 w/cm2), where the liquid flow and the sliding of individual bubbles prevented the
vapor fmm agglomerating above the heater surface.
For flow rewetting experiments under reduced gravity conditions, only a few have
been perfonned. Kawaji et al. (199 1) performed experiments on rewetting of a transparent,
horizontally oriented, 14 mm LD. quartz tube by injecting subcooled R113 under reduced
gravity conditions aboard the NASA's KC-135. Inverted annular Rows with thick vapor films
and dispersed flows with liquid filaments and droplets were observed during the rewetting
tests. Westbye et al. (1995) performed sirnilar rewetting experiments using R-113 and a thin-
wailed stainless steel tube, with 1 1.3 mm LD. and 9 14 mm length aboard the KC-135. The
rewetting temperature and heat transfer coefficient in the film boiling regirne were
significantly reduced due to thicker vapor films under reduced gravity. Quenching curves
obtained under reduced gravity were very sirnilar to those obtained in 1-g, but were shifted to
Iower wall superheats which led to the conclusion that transition and nucleate boiling regimes
are relatively unaffected by the gravity level. Also the critical heat flux was observed to
increase with increasing flow rate in both normal and reduced gravity.
Very recently, Antar and Collins (1997) claimed that a new vaporlliquid flow pattern,
called the filamentary flow pattern, exists under the low gravity conditions after they
conducted quenching experiments involving a quartz tube and iiquid nitrogen aboard the KC-
135. This new flow pattern, which also resembled those observed by Westbye et al. (1991),
comprised of long and connected liquid columns surrounded by a thick vapor layer and
flowing in the center of the test section. The vapor layer was found to be much thicker than
that occurring in the inverted annular flow pattern observed on the ground. Similar to
Westbye et al.'s (1995) results, they dso obtained lower quench speeds in vertical quenching
of a stainless steel tube using liquid nitrogen under low gravity conditions.
In a different type of experiment from pool boiling, flow boiling or quenching
experiments under reduced gravity conditions, Qiao and Chandra (1995) photographicdly
observed boiling of single droplets of water and n-heptane impacting a hot stainless steel
surface during 55 ms free fall. Their study mainly focused on the measurement of Leidenfrost
temperature. They found that the Leidenfrost temperature could not be defined on the basis of
an evaporation curve because droplets did not stay at the surface during film boiling in low
gravity. Due to the effects of surface thermal properties, surface tension and wettability, and
the extent of droplet break up and recoalescence, the behavior of water was different from n-
heptane.
In a theoretical work related to the rewetting mechanism, Adharn-Khodaparast et ai.
(1 994% 1994b) and Adham-Khodaparast ( 1995) studied the Rayleigh-Taylor and Kelvin-
Helmholtz instability of a liquid-vapor interface under the influence of adverse gravity field
and with appreciable interfacial heat and mass transfer. They developed a new mode1
incorporating the effect of vapor recoil that correctly predicts parametric dependence of
critical heat flux reached during rewetting of a hot surface. Their snidy showed that
horizontal velocity difference between the vapor and liquid smams increased the heat flux
requirement for interfacial stability and the effect of gravity became smail if there was
horizontal relative motion between the vapor and liquid in film boiling.
1.3. Research Objectives
It can be summarized from the literature review presented in the previous section that
some pool and flow boiling heat transfer studies in reduced gravity have been made.
However, they are still too few to fully clarify certain issues, such as the effect of gravity
level on nucleate and transition boiling heat transfer, maximum heat flux, rewetting
temperature and quench velocity under different flow and liquid subcooling conditions. With
the exponential growth of space-related activities, such as the use of satellites in
communication, launching of the International Space Station, as well as Mars and Lunar
exploration, the boiling heat transfer data are urgently needed for power and thermal
management systems as well as liquid propellant and cryogenic fluid handling in space.
Therefore, the objectives of this study were to design and build a test loop and test section,
and conduct experiments both on the ground and under reduced gravity to study the effect of
gravity level, liquid flow rate and inlet subcooling on flow boiling heat transfer during
quenching of a hot surface. The main feature of the new expenmental apparatus was the use
of a recently developed micro-sensor for surface heat flux and temperature measurements.
In addition to extensive tests on the ground, a senes of expenments have been
performed aboard the KC-135 and DC-9 parabolic aircraft operated by the U.S. National
Aeronautics and Space Administration (NASA) in order to obtain data under reduced gravity
conditions. In the rest of this thesis, a detailed description of the experimentai apparatus will
be presented in Chapter 2 and the characteristics of reduced gravity condition aboard the
aircraft as well as the experimental procedures will be descrîbed in Chapter 3. Then the
experimental results, and data andysis will be presented in Chapters 4 to 8. Finally, Chapter 9
presents the concIusions and recornmendations for future studies.
Experimental Apparatus
Based on the previous KC-135 campaign experiences gained by two other
researchers, Westbye (1992) and Lui (1993), a new experimental apparatus was designed and
constnicted to collect rewetting data aboard the NASA's KC-135 and DC-9 parabolic
aircraft. Several criteria were considered in the new design. The whole system was designed
to be as light and compact as possible to make operation, shipping by air freight and loading
ont0 the KC-135 and DC-9 more reliable and simpler. Two compact and portable modules
were used to contain a test section, flow loop and heat transfer system in one and a data
acquisition system in the other. The 80w loop was designed to provide fluid flow, boiiing
heat transfer, cooling and phase separation even under microgravity conditions with R113,
PF5060 (Per-Fluorocarbon, C6FL47, a new refrigerant from 3M) and other coolants for the
present and future flow boiling experiments. The fluid properties will be given in Chapter 8.
At the beginning of flow boiling experiments in this study, R113 was used as a
coolant because of its non-toxic and non-flarnmable nature and especially its low boiling
point. However, this fluid has been recently banned due to its harmful effect on ozone layer
of the atmosphere. Thus, a new substitute, PF5060, a coolant with sirnilar therrnophysical
properties as R113's, was found to continue this study. The test section was constructed to
enable surface heat flux and temperature measurements for the present rewetting
experiments. The data acquisition system was set up to monitor and collect fiow and heat
transfer data automatically. The operation procedure was organized to be as simple as
possible, so that one operator could perform the entire experiment aboard the KC-135 and
DC-9 aircrafts. A general description of the apparatus is given in the following sections.
2.1. Test Loop and Components
The rewetting experiments had to be conducted aboard KC-135 and DC-9 parabolic
aircraft of NASA. Therefore, the expenmental apparatus was designed for operation under
reduced gravity conditions and to withstand harsh crash conditions (9-g fonvard 2-g upward
and 7-g downward) aboard the aircraft in cornpliance with the requirements from NASA.
The test loop schematic is shown in Figure 2.1. The working liquid (R113 or PF5060)
was circulated through the closed flow loop by a speed-controlled Micropump (Model 10 1-
415), with a maximum flow rate of 26 Ipm and a head of 200 kPa. This pump was driven by a
motor with a magnetic coupling that did not require any sealing mechanism or lubrication.
The flow rate was measured by an infrared flow sensor (IR-OP Flow Meter Model 502-104)
and KEPtrol LED-display unit with an output signal of 0-5 V. The accuracy of flow
measurements determined by calibration was M.01 (literhinute). Two T-type thermo-
couples and two Omega PX176-025A5V strain gauge pressure transducen were used to
measure the inlet and outlet temperatures and pressures of the test section, respectively.
While passing through the test section, the working liquid was partially vaporized and the
resulting two-phase mixture entered the condenser, where the vapor was condensed into
liquid and the liquid was further subcooled.
Cornputer \ e Relief
Test Valve section T
amcorder l - œ - - Condenser 1 - - Oi O ooler 1 EXP-16 Camcorder - Micro
.IIIILIIIIILI
I I I s e x t u 3 2 a - & L - Discharge Pump
Line e
iT: E S \
s!
Figure 2.1. Schematic of the test loop.
Both fluids, R 1 13 and PF5060, were cleanly handled in the flow loop using the
chernically compatible Teflon tubing andor brass fittings to connect various flow loop
cornponents.
The experiments under reduced gravity were performed aboard the NASA's KC-135
and DC-9 aircraft in Houston and Cleveland, respectively. The experimental apparatus was
built inside two aluminum frame assemblies measuring 0.508m x 0.508m x 0.787m and
0.7 1 lm x 0.508m x 0.9 14m. The components of the test loop and the data acquisition system
were mounted inside the multi-level frarne assemblies that could easily be loaded ont0 the
aircraft and fastened to the cabin floor.
2.2. Heat Flux Micro Sensor
The main god of the study was to obtain instantaneous surface heat flux and
temperature data dunng rewetting of a hot surface so that the mechanism of boiling at high
surface temperature could be better understood. Previous experiments dealing with boiling
and rewetting studies often used thermocouples for temperature measurement. Usually, the
temperatures were measured below the heated surface of the test section and not on the
surface itself. Even if surface thermocouples or resistance temperature devices were used,
there were no direct means of heat flux measurement. In the previous studies, heat flux was
always indirectly calculated from the thennocouple measurements using an inverse
conduction computational technique. In most of those studies, even the surface temperature
was to be calculated from the inverse conduction calculations. This was a major source of
error and uncertainty in those experiments, because many transient and dynamic features of
the heat transfer problem could have gone undetected in those experiments.
In the present study, the Heat Flux Micro Sensor (HFMS) of VATEIL Co. was, for
the first time, used to simultaneously and directly rneasure the surface temperature and heat
flux variations during rewetting. The HFMS consisted of two heat flux gages having a square
shape and 2rnm by 2mn: in size, and placed 7 mm apart (center to center) with a platinum
Resistance Temperature Sensor (RTS) in between, as shown in Figure 2.2. The sensors were
deposited on a copper or stainless steel circular disk with a plasma-sprayed, 100-micron thick
aiuminum oxide base insulating layer. Also, the sensor had an over-coat of sputtered 1
micron thick aluminum oxide that provided moderate abrasion resistance and electncal
isolation. The sensor was developed by the Vatel1 Corp. using the direct rnetal deposition
technique from the electronics industry, in the form of micron thick thermopiles of
platinum/platinum- 1 0%-rhodium (Hager et al. 1989).
The thermopiles were layers of very small and thin themocouples which were
altemately deposited above and below a very thin thermal insulation layer (as illustrated in
Figure 2.3). For any finite heat flux through the thermopile, there would be a temperature
difference between the upper and lower faces of the insulation layer. Therefore, the
thermocouples above the plate would sense a temperature difference and produce a certain
electrical motive force (emf). However, the subsequent upper and lower face thermocouples
were connected in a reverse senes, so that the positive junction of one upper thermocouple
was connected to the positive junction of one lower thermocouple and the same for the
negative junctions. This arrangement gives a zero emf for the case of no heat flux. On the
other hand, for any non-zero surface heat flux, this combination of lower and upper
themocouples gives a small electrical potential output equal to the difference between the
emf's of each thermocouple. Severai sets of the upper and lower themocouples were
connected in a reverse series to give a measurable output signal for the thermopile.
The output of the HFMS was linearly dependent on heat flux, because the thermal
resistance of the insulation layer was constant and temperature difference across the plate was
linearly dependent on heat flux. The output of each KFMS had been calibrated by the
manufacturer and was 2.2 micro-voit for a heat flux of 1.0 w/cm2. Although the sensors are
resistant to temperatures as high as 1,200 O C , they are very delicate and thin, therefore, care
was needed to avoid mechanical damages during assembling or disassembling the test
section.
The sensor connections were made of platinum pins with spot-welded platinum leads.
Ordinary copper wires could be used for connection to those platinum pins as long as there is
no temperature difference between them. In order to reduce the noise as much as possible, a
shielded, twisted pair of wiring was used, and its total length was kept to a minimum.
Micro Heat Flux Sensors
Resistance Temperature Sensor
Figure 2.2. Schematic of the heat flux micro sensor.
Lower U P W Thermocouple Therm ocouple
,ead inection
u I I
: n units of thermopiles
Figure 2.3. It shows the working principle of the heat flux micro sensor (Hager et ai. 1989).
Two dfierential amplifien from ECTRON Co. were used for amplmng the output
signals from the heat flux gages. They were ECTRON mode1 687 rnounted in a bin that used
115 VAC or 12 VDC power (12 VDC power was used in this study to avoid any electrical
noise). The differential amplifier ensures the best low noise performance of the heat flux
sensor since it has transformer isolation between the signal inputs, power supply and output.
The amplifier gain could be adjusted in steps of 10, 20, 50, 100, 200, 500 and 1,000 with a
0.5% accuracy. A twenty-turn front panel potentiometer provided variable gain from
approximately 0.95 to 2.7 times the switch selected value. In this work, a gain setting of 1 0
was used in al1 experiments. The uncertainty in the heat flux measurement was analyzed and
estimated to be fi1 of the reading, which is presented in Appendix II.
The instantaneous surface ternperature was measured by the resistance temperature
sensor (RTS). The electncal resistance of the RTS changes linearly with temperature as
indicated by the calibration data supplied by the manufacturer. It operates on the principle
that the electrical resistance of a metal changes when subjected to a change in ternperature.
The RTS was excited by a precision DC voltage supply and formed one of the amis of a
Wheatstone bridge configuration. The excitation current was adjusted with regard to the
range of resistance of RTS, so that it would not produce appreciable self-heating in RTS. The
typical output of the sensor was about 300 micro-volts per 10 O C at 24 OC. The uncertainty in
the surface temperature measurement was estimated to be rr2 OC as presented in Appendix II.
2.3. Test Section
The test section consisted of a 19.1 mm thick stainless steel base plate and 6.4 mm
thick alurninum casing which formed a rectangular flow channe140 mm wide by 5 mm high
and 200 mm in Iength. In the present experiments, the boiling heat transfer characteristics at
high wall superheats were studied using a quenching method. Therefore, only one of the
horizontal surfaces of the test section was electrically heated.
The schematics of the heated stainless steel plate housing one HFMS, two cartridge
heaters and four T-type thermocouples are shown in Figure 2.4. The plate was 200 mm long
and 50 mm wide with a 5 mm groove around the plate penmeter to house a Teflon seal,
which couId effectively prevent the leakage of R113/PF5060 without melting at the wdl
temperature of up to 280 OC.
r--> A I
3 , , 1 L--> I B \
Cartridge stainléss Aluminum Flow HFMS Window HFMÇ Heater Steel Base Casing Channel Housing Glass Wire Hole \ \ / / / /
Section AB
Figure 2.4. Schematic of the test section.
The HFMS was imbedded in the plate at a distance of 50 mm from the inlet and was
flush with the heated surface. A brass tube was used to fasten the HFMS to the plate. The tip
of the tube was rnachined to provide a 3 mm screw that fitted into the rear hole of the HFMS.
The outside perimeter was threaded and a bolt was used to pull the tube and HFMS towards
the bottom of the plate. A Teflon gasket was used to seal the bottom of the HFMS against the
edge of its housing.
Four shielded thermocouples with bare tips were mechanicdly fixed in holes dnlled
to 1 mm beneath the rewetting surface. They were used to monitor the temperature
distribution dong the flow direction and check its uniformity before each experiment. The
thermocouples had an accuracy of fl. 1 O C .
The stainless steel plate was heated by two 600W Watlow Firerod LA Cartridge
heaten which were connected to an AC transformer for variable power input. Two cartridge
heaters were screwed into two 8 mm holes parallel to the plate length. The extemai surface
of the plate was fully insulated by fiber glass.
The aluminum casing of the test section. which provided the other three sides of the
flow channel besides the heated wall is shown in Figure 2.5. There were three glass windows
aaached to enable direct viewing of the flow and boiling conditions above the HFMS. The
casing was machined out of a 19 mm alurninum plate with two grooves, one matching with
the heated wail groove and the Teflon seal, and the other with the flow channel. On the upper
face, a circular window was provided by machining a 25 mm hole with a 1.5 mm inside
shoulder to provide support for mounting and Epoxy gluing of a circular 28 mm diarneter
Boro-Silicate g las piece. This circular window was situated directly over the HFMS. Two
other windows were each situated on the side wall of the casing at a distance of 50 mm from
the inlet. On each wall, two rectangular openings were machined with dimensions of 23 mm
x 5 mm and 30 mm x 10 mm. On the outer and Iarger opening a glass piece of size 30mm x
10 mm was fixed by high temperature Epoxy glue and on the inner opening was mounted a
glass piece of size 23 mm x 5 mm. The outer glass provided the mechanical strength of the
window and the inner glas was installed flush with the inner wall so that no disturbance
would be introduced into the flow inside the channel.
Flow Casting
Circular Window
w /
Flow Passage d
ined tube
Flow
/
Cylindrical O pening Conical Opening
\ /
Figure 2.5. Schematic of the aluminum case cover.
Figure 2.6 shows the whole view of the test section after assembly. The aluminum
casing and the heated wall were connected and fastened by four aluminum clamps. The idet
and outlet sections were screw fastened to the flow channel and provided a smooth passage
between the channel and round tubes. The inlet and outlet sections were designed with the
same geometry so that flow could be directed in both directions through the test section. They
were machined out of 19 mm thick aluminum slabs. A 15 mm inner diameter hose connector
was machined out of aluminum and connected to a rectangular collection box with
dimensions of 65 mm x 22 mm. The rectangular box was machined inside to provide a
conical gathering section leading to a cylindrical outlet, concentric with the hose connector.
This design with its smooth contraction and expansion provided least resistance to the two-
phase flow entenng or leaving the test section.
Teflon gaskets were used for sealing al1 the components of the test section with the
exception of the glass windows where high temperature Epoxy was used.
2.4. Condenser
It is very important in flow boiling experiments that liquid can be continuously
supplied to the test section, free of any vapor. This is a trivial problem for the experiments on
the ground under 1-g conditions, since the large density difference between vapor and liquid
causes the two fluids to separate naturally. However, because the surface tension is the only
dominant force under reduced gravity, liquid-vapor separation in condenser becomes a major
technical problem in order to continuously supply liquid to the test section. Therefore, Iiquid-
vapor separation must be considered carefully in design of any flow boiling experiments
under reduced gravity.
Two major methods can be used in the design of a liquid-vapor separator for reduced
gravity flow boiling experiments. One is to introduce an acceleration field so that liquid and
vapor can be separated based on a density difference. This cm be achieved by rotating the
separator to form a centrifuga1 force field. However, this kind of separator would be
complicated in design and expensive to make, and it would only be needed to provide high
flow rates of liquid during long reduced gravity periods. The other option was the utilization
of liquid-surface wettability. The excellent wettability of Lucite, especially in the case of
highly wetting liquids like R-113 and PF5060, can be used for liquid-vapor separation. The
present design of condenser in this study utilized the application of this method. It could
provide a relatively simple means of liquid-vapor separation for short duration of reduced
gravity experiments.
A compact condenser, which was especially designed for KC-135/DC-9 reduced
gravity experiments, consisted of a transparent cylindrical vessel, 203 mm in inner diameter
and 406 mm in height, made from 6.4 mm thick Lucite. Donut-shaped Lucite plates were
glued to the cylindrical shell to act as flanges for the top and bottom covers. The vessel was
closed by two circula Lucite plates which were fianged to the top and the bonom of the
cylindrical shell with steel bolts and nuts. Rubber ring-shaped gaskets wrapped by Teflon
tape were used to seal the joints. A pressure gauge was installed on the top of the condenser
to monitor the inside pressure. The interna1 components included an inlet spray head, copper
cooling coil, a cap, a bubble ejection (or liquid collection) device, wire mesh and a cooling
system as shown in Figure 2.7. The transparent vessel perrnitted a video carncorder to view
and record the fluid motion in the condenser during reduced gravity experiments.
Two types of inlet noules attached at the top center or near the side wall, were
designed to force the incoming vapor-liquid R-113 mixture to flow downward as a spray
cone or in a spiral dong the vesse1 wall. This design was expected to not only enhance
condensation heat transfer, but to continuously supply liquid R-113 to the top of the liquid
collection device even under reduced gravity so that the condenser outlet would rernain
irnrnersed in liquid during both reduced gravity and hyper-gravity periods. In practice, the
spiral inlet gave a better result in reduced gravity experiments. A combination of a bed of
wire mesh, a perforated copper cap and bubble ejection fins was used to prevent the vapor
from rnoving towards the condenser outlet Iocated at the bonom of the condenser through the
effect of surface tension. The design was based on the tendency of liquid to obstmct the
passage of vapor through narrow channels and contracting conduits. The bubble ejection fins
(or altemately called liquid collection device) consisted of 24 pieces of 7.5 cm long, 12.5 cm
wide and 1.6 mm thick Lucite plates joined by two Lucite connectors and providing 15'
radially contracting passages for the liquid to flow towards the pump suction. The two-phase
mixture was introduced at the top of the separator and any fluid traveling towards the pump
Cooling Coil .
W ire, Mesh
Figure 2.7. Schematic of the condenser.
suction had to go through the condenser section contining copper cooling coils where most
of the vapor was condensed, then the wire mesh bed, the perforated copper cap and through
the bubble ejection fins.
The cooling system included a March MDX-3 centrifugai pump, a 10" Electra-fan, a
B&M Supercooler heat exchanger, cooling coils and water as the coolant. Two cooling coils
made of 6.4 mm O.D. copper tubes were placed inside the condenser to condense the vapor
and cool the liquid. The center coi1 was finned to enhance heat transfer. To avoid king
classified as a pressure vessel, a pressure gauge, a relief valve and a filling valve were
installed at the top of the condenser.
2.5. Data Acquisition System
A Perfect 486DX-33Pentium- 100 persona1 cornputer was used to gather and store ail
the expenmental data including the heat flux and surface temperature, heated plate
temperatures, inlet and outlet fluid temperatures, R-113lPF5060 flow rate and gravitational
acceleration level. Data were collected via Keithley Metrabyte EXP- 16 16-channel expansion
mutiplexer, signal conditioning board and DAS-1402 data acquisition board capable of up to
100,000 samples/sec sampling rate. Figure 2.8 shows the data acquisition flow chart. Labtech
Notebook software was used for data acquisition management allowing real-time
visualization of heat flux and surface temperature histories, which was very helpful to the
operator for producing smooth initial temperature profiles and obtaining better quenching
data aboard the KC-135fDC-9 aircraft. The data were analyzed using Iandel SigmaPlot
software.
3 Gravity Levels >i
2 Pressure
Figure 2.8. The flow chart of the data acquisition system.
, > Personal
Cornputer
Transducers
Flow Meter
2.6. Visualization of Quenching Experiments
2.6.1. Quenching of a hot surface with subcooled RI13
A boiling process is commonly encountered wherever a heating surface is at a certain
temperature above the boiling point of a Iiquid. However, the mechanism of boiling is not yet
fully understood. Visurlization is an important tool in two-phase 80w and boiling heat
transfer studies, since it can efficiendy help researchers investigate the flow and boiling
mechanisms. In order to further analyze the flow boiling phenomena investigated with the
advanced heat flux micro sensor in the present study, the test loop and the test section
descnbed in the previous sections were used for the visualization of quenching expenments
(RI13 was used as coolant) by using a video carnera and recording system. Two sets of
expenments were conducted.
Since the total size of the experimental apparatus aboard the KC-135 aircraft was
limited, an ELMO micro video carnera (mode1 EM-102BW) and a Mitsubishi carncorder
Amplifier #2 X I 0
HIF Sensor #2
6 T-Type Themocounles -
> + +
- EXP-16 X200 RTS
+
+ I >
Bridge >
DAS 1402 A/D 1 Ml KHz
XI 1 r
EXP-16
>
were used to record the boiling phenornena occumng on the micro heat flux sensor surface
through the top glass window of the test section. The frame rate of the video recording
system was 30 Hz.
In the ground tests, a high speed video camera was used to record the quenching
process via a 45' inclined mirror placed above the top glass window of the test section and
located above the heat flux sensor. At the sarne time, it aiso recorded the heat flux level by a
heat flux indicator placed beside the mirror. The heat flux indicator consisted of a ten-level
indication LCD and ten glas fibers (1 mm in diameter), as shown in Figure 2.9. The outlet
signal from the heat flux micro sensor was amplified and sent to the heat flux indicator while
it was recorded by the data acquisition system at the same time. One side of glas fibers was
inserted into the LCDs and fixed by Silicon rubber. The other side was aligned on a piece of
plate. When the input signai was above a certain threshold value of an L m circuit, a certain
number of LCDs proportional to the input signal would light on. The glass fibers transferred
the light from LCD's to the plate, which faced the high speed video camera The heat flux
indicator was set to show the heat flux from LOO kw/m2 to 1000 kw/m2. The accuracy of the
indicator was determined by the threshold value of the LCD circuit and was about 3 0
kw/m2.
2.6.2. Quenching of a hot rectangular quartz tube with subcooled water
The test loop consisted of a pump with a fixed flow rate, a rectangular quartz tube as a
test section, a bypass flow loop, a tank made of Lucite plates, three valves, transparent plastic
tubing and brass fittings, as shown in Figure 2.10. The quartz tube was 10 mm by 6 mm in
cross section and 250 mm in length and was placed vertically. The wall temperature was
measured by a T-type thermocouple fixed on the tube by a piece of copper tape. At the
beginning of the experiment, the test section valve was switched off so that the distilled water
was circulated in the bypass flow loop. After one side of the quartz tube wall was heated by a
propane torch to 500 - 600 OC, the test section valve was opened to a designed position
corresponding to a small, intermediate or high flow rate. Then the distilled water was
circulated to the test section and was partially evaporated while quenching the hot wall of the
quartz test section. The quenching process was recorded by a high-speed video camera at a
frame rate of 744 Hz and a shutter speed of 1110000 second.
Test Section
HVC - I 1 Tank
>
i
Memory Bypass Valve 1
Test Section I 1 Valve
Pump
Discharge Line
Figure 2.10. The schematic of the test loop in quenching of a quartz tube with water.
CHAPTER 3
Reduced Gravity Experiments and Procedure
The descriptions of KC- 135/DC-9 aireraft and rewetting experiments are presented in
this chapter. Reduced gravity experirnents were performed aboard the KC-135 and DC-9
aireraft dunng the penods February 1-4, 1994 and March 4-7,1997 at NASA's Johnson Space
Center and Lewis Researeh Center, respectively. A large arnount of data has k e n collected
and processed using the methods of data analysis to be described in Chapter 4.
3.1. KC-135 and DC-9 Aircraft
There are four ways of conducting reduced gravity expenments. Drop towers which
can provide reduced gravity levels in the order of lodg (g denotes the terrestrial gravity, 9.8
m/s2) conditions at intermediate cos& are cornmonly used for the experiments which do not
require long duration (not more than 5 - 10 seconds) of reduced gravity. Sounding rockets
can provide a few minutes of high quality zero gravity (in the order of 1 0 3 ) at a
considerably higher cost. Satisfactory rnicrogravity environment cm be established on the
Space Shuttle and space stations, but the availability of fiight opportunities is very limited
and the costs are also very high. Parabolic flying aircraft such as the NASA's KC- 135 and
DC-9 can provide a series of about 15 to 20-second long reduced gravity periods dunng the
flight (in the order of t 1 0 * ' ~ ) and are ideal for the expenments which do not require precisely
zero-gravit- conditions.
In this work. a platfom for performing the rewetting expenments provided aboard the
KC-135 and DC-9 was utilized, because each rewetting expenment for the present test
section could be finished in about 15 - 20 seconds.
The KC-135 aircraft is a four engine swept wing aircraft similar to a Boeing 707, and
has been extensively modified to support reduced gravity experimental research. It is
operated by the NASA Johnson Space Center (JSC) Reduced Gravity Prograrn in Houston,
Texas. Penods of reduced gravity and hyper-gravity are provided by flying a parabolic arc
between the altitudes of 24.000 ft (7.3 km) and 32,000 ft (9.8 km). From a level fiight, the
aircraft is fmt orientated into a " 45' nose-up" attitude. The engines are then throttied b ~ k ,
until thmst equals drag, and the KC-135 coasts over the crest of a parabolic arc. Thus, the
period of reduced gravity is sustained for about 20 seconds, and sometimes up to 30 seconds
with a typical acceleration within fl% of normal gravity (g). The parabolic manoeuvre is
initiated and terminated with a pull-up and pull-out of about 1.8 g acceleration. Each cycle of
reduced gravity and hyper-gravity lasts for about 70 seconds.
Four sets of 10 parabolas are flown dunng each flight, and a five to ten minute period
of level flight is provided to allow modification of equiprnent, replacement of video tape.
checking of the data recorded by data acquisition system, and so on. Typical acceleration
levels in X-Y-Z directions during reduced gravity periods aboard KC-135 are shown in
Figure 3.1.
The performance of DC-9 is similar to KC-135's. The reduced gravity period
provided by DC-9 usually lasts for about 15 seconds. In the DC-9 campaign, there were about
50 parabolas flown in each Right. Because the flying region is limited to an area above Lake
Huron, the number of each set of parabolas could not be pre-specified. In the DC-9 campaign,
only z-acceleration (perpendicular to the cabin floor) was recorded.
3.2. Operational Procedure
The following procedure was used during the KC-135 campaign in February 1994.
Preparation before flight:
Degas the air and non-condensable gases out of the liquid in the test loop and
condenser by circulating and boiling R 1 13 after the loop is filled with R 1 13.
Circulate RI 13 and tum on the power to the test section's cmridge heaters and
pre-heat the liquid at full power to a specified subcooling. Before starting
parabolic flights, heat the test section to a desired temperature.
During hyper-gravity:
0 Tum on the heater at full power. Since it requires one to two minutes at full
power to heat the test section from about 50 OC to a suitable initial temperature
of about 200 OC, skip altemate parabolas to allow the test section to heat up.
Time (second)
O 5 10 15 20 25 30 35
Time (second)
Time (second)
Figure 3.1. Typical accelerations during the KC- 135 flights
During reduced gravity:
* Tum off the heater power. Start to record the data. hject the liquid into the
test section at a specified flow rate. Continue recording the data after
quenching.
This procedure was repeated for each parabola. For DC-9 campaign and normal
gravity tests on the ground, a similar procedure was used. Data collected by the data
acquisition system during each run were saved to a file, and the system was reinitialized.
3.3. Condenser Performance and Flow Rates
The condenser which was especially designed for the KC-135 reduced gravity
carnpaign could successfully supply enough liquid to the test section at almost al1 flow rates.
The wire mesh effectively prevented the liquid contained in the reservoir section (lower part
of the condenser) from nsing upward during reduced gravity. In the KC-135 carnpaign, the
wire mesh was 10 cm thick and the initial liquid level was almost at the sarne level as the
upper part of the mesh.
Figure 3.2 shows the reservoir full of liquid under normal gravity or hyper-gravity
conditions. Under reduced gravity and at low flow rates, the liquid compietely filled the space
around the bonom outlet of the reservoir. The bubble columns entering the reservoir from the
side wall below the condenser section, were slowly moving downward but did not reach the
bottom of the reservoir. They were effectively blocked by the liquid collection device to stay
around the outer edge of the liquid collection device (see Figures 3.3 and 3.4). At
intermediate flow rates, the bubble columns reached the bottom of the reservoir but did not
collapse towards the center as shown in Figure 3.5 so that no bubbles entered into the pump
suction and the flow loop. Figures 3.6 and 3.7 show that at higher liquid flow rates, the liquid
collection device sometimes failed to completely prevent the bubbles from reaching the
bottom center. This was improved in the DC-9 campaign by closing the top of the liquid
collection device and leaving the side wall as the only way for the liquid to refill the
reservoir.
In the KC-135 campaign, the reservoir below the condenser in some runs could not be
continuously refilled with the liquid because the wire mesh could not efficiently capture the
liquid lying above the mesh and the inlet nozzles were located too high above the mesh. This
was improved by increasing the height of the mesh and placing the nozzles close to the mesh
in the DC-9 campaign.
Row rates were very hard to control smoothly and accurately at the begiming of each
rewetting experiment under reduced gravity due to the operator's action being affected by the
reduced gravity condition. The typical flow rate profile in KC-135 campaign is shown in
Figure 3.8 for one run, which also presents the corresponding z-acceleration level in reduced
gravity experiments. The acceleration level normal to the heated surface ranged between
rr0.02 g. At the beginning there was a short transition petiod fiom no flow to a desîred flow
rate. This duration was less than 1 second for some runs but about 2 to 4 seconds for other
runs. The effect of this delay could be detected in the heat flux and surface temperature
histones. During the quenching period the liquid flow rate experienced only small changes
which would have linle effect on the quenching process. These effects have been considered
in the anaiysis of the experimental data presented in the next chapter. Figure 3.9 shows the
data collected during the DC-9 campaign. The drop in flow rate between 14 and 15 seconds
was caused by a vapor bubble entering the test section. The z-acceleration level could be kept
between kû.02 g for about 15 seconds.
It is apparent from the z-acceleration levels shown in Figures 3.8 and 3.9 that the
frequency of g-jitter aboard the DC-9 was much lower than that on the KC-135. The vibration
produced by this low fiequency g-jitter caused the wire mesh to easily fail to prevent the
liquid in the reservoir frorn floating upward and the vapor column from entering the reservoir
even when there was no flow in the test loop. This made it more difficult to conduct the
experiments at lower flow rates. The performance of the condenser was not recorded by a
video camcorder in the DC-9 campaign.
The frequency spectra of g-jitter in the KC-135 and the DC-9 were very different.
However. the effect of such g-jitter on quenching characteristics and boiling heat transfer was
found to be smail in the present work. The detailed analysis is presented in Appendix I(2).
Figure 3.2. The reservoir full of Iiquid in normal gravity and hyper-gravity.
Figure 3.3. The blocked bubble columns at low fiow rates in p.-g.
Figure 3.4. The blocked bubble columns at low flow rates in p-g.
Figure 3.5. The blocked bubble columns at intermediate flow rates in p-g.
39
Figure 3.6. The blocked bubble columns at high flow rates in p-g.
Figure 3.7. The collapsed bubbles entering the flow loop in p-g.
40
6 0.08 RUN MG423 -
Flow Rate 5 - - 0.06
Time (sec)
Figure 3.8. Typical flow rate profile for one run in p-g (KC- 135).
Figure 3.9. Typical flow rate profile for one run in p-g (DC-9).
41
3.4. Quenching Experiments
Two flights of forty parabolas each were avaiiable to perform quenching experiments
during the KC-135 campaign. The quenching tests were performed by preheating the plate to
about 220 OC and injecting liquid RI 13 to quench the hot surface. Considering the safety
problems and the time required for full quenching and the penod of reduced gravity available,
the plate temperature in reduced gravity experiments was set to a value ranging from 130 OC
to 222 OC. Lower initial plate temperature could be reached between two parabolas, but a
higher temperature needed longer heating periods due to the limited electricd power
available aboard the KC-135. Therefore, a few parabolas had to be skipped and this lirnited
the number of runs that could be performed in a given flight.
The inlet liquid temperature increased gradually from 7 OC to 22 OC dunng the first
flight and from 22 OC to 41 OC in the second flight, because the cooling system could not
remove al1 the heat generated by the test section in a very short time between two or three
parabolas. Since each quenching experiment lasted for about 20 seconds, the inlet liquid
temperature did not increase significantly dunng each run.
The liquid flow rate ranged from 1.5 literlmin (ji = 0.125 d s ) to 1 1 literlmin (jI = 0.92
mis), but it was dificult to repeat the same flow rate. In addition to the reasons mentioned
above, the pump head and flow rate changed with the back pressure in the test loop, which
increased during quenching due to the generation of a large arnount of vapor in the test
section. This contributed to the difficulties in controlling the flow rate.
Since there were several problems beyond the operator controI under reduced gravity
conditions such as unstable flow rate caused by bubbles entenng the test section and larger
than expected amplitudes of fluctuations in acceleration levels, a total of 12 reduced gravity
mns, 8 negative reduced gravity runs and 16 ground runs were used for analysis as listed
below in Table 3.1.
Unlike in the KC-135 carnpaign, PF-5060 was used as a coolant in the DC-9
campaign. Its rewetting temperature and quenching speed are lower than R- 1 1 3 ' S. In addition,
the penod of microgravity lasted for about 15 seconds, which is shorter than KC-135's.
Therefore, the suitable setting of initial temperature was crucial to ensure the full quenching
Table 3.1. The List of p-G and 1-G Runs (KC- 135 campaign).
Reduced Gravity Normal Gravity
MG-D2- 133- 1.4- 1 1 MG-DO- 1464.0-7
Subcooling - 25 O C
NG-G3-235- 1.5-23 - -
MG-D 1-1 30-4.8-8 MG-D23-207-4.4-22 MG-D8- 180-4.9- 13 MG-D12-177-5.9-16 MG-D25-222-6.8-2 1 MG-D22-205-5 -5-2 i
NG-G4-247- 1.7-2 1 NG-G2-237-2.5-2 1 NG-GO-220-3 -3-20 NG-G 1-222-3.5-22 NG-G5-248-4.8-24 NG-G9-253- 1 1.4-2 1
-
MG-D 1 9-205-9.2-2 1 MG-DSO-2 13- 10.0-2 1 MG-D 15- 192- 10.4-2 1 MG-D2 1 - 194- 10.6-22
1 NRG-E 15-203- 1.6-24 1 Subcooling < 4 O C 1
NG-G 10-288- 1 1-1-26
Subcooling - 16 OC A
NG-G8-257- 1.6-32 Negative Reduced Gravity NRG-E30- 160- 1.1 -4 1 NRG-E9- 165- 1.1 -24
I NRG-E 18- 168-5.3-29 I NG-GS2-200-4.0-44 I
NG-G6-248-6-2-30 NG-G7-249- 1 1.3-32
Note: In the table above, the nui number is designated to indicate the experimental
conditions as follows.
For example,
MG - 023 - 207 4.4 - 22
Reduced gravity - nui number - initial wall temp.("C) - f iw rate(hin) - inlet temp.("C)
and
NRG - negarive reduced gravity (-0.1 g f0.02 g )
NG - n o m 1 gravis
D, E, G, GS - represent particular date of expehents
of the HFMS surface in 15 seconds. After one flight trial, the suitable initial temperature was
found to be about 150 to 160 OC. The inlet liquid temperature changed from 20 to 34 OC in
all the DC-9 flight tests. Because of the effect of low Frequency g-jitter, the flow rate ranged
from 5 litedmin (il = 0.42 m/s) to 12 litedmin (ji = 1 mls) in this campaign. In summary, a
total of 17 reduced gravity runs and 16 ground runs were analyzed, as listed in Table 3.2.
3.5. Heat Flux and RTS Sensor Performance
It is difficult to calibrate the heat flux micro sensor ( H F M S ) over a large measurement
range due to the unavailability of the required instruments or facilities. The manufacturer
performed calibration of the heat flux micro sensor over a short range of heat flux (450
kw/m2 maximum) by using free jet convection. The sensitivity of the sensor to temperature
was estimated to be about fi% of the reading (shown in Figure 3.10). As described in
Chapter 2, the output signal from each heat flux sensor was amplified by a pre-amplifier with
a gain of 100 and a main amplifier with a total amplification factor of 20,000. Considenng
the offset, the following equations were used to convert the input voltage, V in volts, to heat
flux, q in kw/m2, in the data acquisition system.
Temperature (OC)
Figure 3.10. The calibration curve for HFMS provided by the manufacture.
Table 3 m 2 m The List of p-G and 1-G Runs OC-9 carnpaign). -- - - -
Reduced Gravity
MG-T720- 1 54-6.7-22 1 NG-P 1 1 - 193-5.2-22 1
Normal Gravity
MG-T72 1 - 166-5.3-22 MG-T69- 153-6-0-23
Subcooling - 33 OC NG-P 14- 192-5.2 1-23
Note: In the table nbove. the run number is designated to indicare the experimenral
conditions as folio ws.
For example,
MG - 1 - 7 2 1 - 166 5.3 - 22
reduced gruviiy - run nwnber - initial wall temp.(T) - flow rate(hin) - Nllet temp. (OC)
and
NG - normal gravity
T. P - represent particular date of experiments
MG-T55- 148-7.0-24 MG-T6 17- 1 54-7 -5-24 MG-T732- 153-7.5-3 1 MG-T6 10- 148-7.7-23 MG-T6 13- 150-8 .O-23 MG-T7 12-15 1-8.5-19 MG-T73 1 - 168-9.0-3 1
NG-P 12- 197-7.2-23 NG-P09- 195-8.2-23 NG-P15- 199-8.3-25 NG-P16- 195-8.8-24 NG-P 13- 195- 10.8-24 NG-P23-209- 12.2-25 NG-P 10-20 1 - 12.7-25
For copper disk HFMS:
q 1 (kw/m2) = 4.37 x 1 04(v - 1.7 x 104)
qz (kw/m2) = 4.88 x 104(v + 1 x 10'~)
For stainless steel disk HFMS:
q 1 (kw/m2) = 3.1 1 x 1 04(v - 1 -84 x 1 o'~) q2 (kw/m2) = 3.56 x lo4(v - 2.26 x L O - ~ )
The resistance temperature sensor (RTS) had been calibrated by the manufacturer
using a unifom temperature oven with a standard thennocouple in contact with the sensor.
Before each set of experiments, the RTS output and the bridge circuit were checked against
the thermocouples used to measure the intemal plate temperatures dong the test section. The
calibration curves for the copper and stainless steel disk sensors are shown in Figures 3.1 1
and 3.12, which show good linearity within t 2 OC frorn 24 to 204 O C and from 21 to 207 OC,
respectively. The following equations were used to convert the sensor output voltage, V, in
volts to wall temperature.
For copper disk HFMS:
Tw (OC) = 2.76 x 104 (V - 4.98 x 10-~)
For stainless steel disk HFMS:
Tw (OC) = 3.78 x 104 (V - 6.70 x 10-~)
In order to capture the frequency of liquid-solid contact, the sampling rate of the
experiments conducted in the DC-9 campaign was set to 500 Hz. This setting enabled
detecting the liquid-solid fiequency up to 100 Hz with up to 5 data points in one cycle. From
previous publications, this sampling frequency was considered to be high enough to
determine the liquid-solid frequency in transition boiling regime as well as in nucleate boiling
regime in the hi& wall superheat region. The pre-amplifier and EXP-16 expansion board
were tested for high frequency cut off by input of a sine wave. The test results showed that
the two pre-amplifiers had a high frequency cut-off at 4 kHz. The high frequency cut-off for
the EXP- 16 board was at 100 W.
250 copper disk
Figure 3.1 1 . The calibration curve for the RTS with copper disk.
Figure 3.12. The calibration curve for the RTS with staidess steel disk.
47
4.1. Transient Heat Flux and Surface Temperature Characteristics The heat flux and surface temperature data were recorded at a sampling rate of LOO
Hz in KC-135 campaign. Figures 4.1 and 4.2 show the histories of transient heat flux and
wall superheat (Tw - TwJ for one of the rewetting experiments conducted in reduced gravity,
respectively. These heat flux and wall superheat transient data are typicai of rewetting
experiments performed under both normal and reduced gravity. Before pumping the liquid
into the test section (from O to 6 second). the heat flux and surface temperature were almost
constant, but there were some high frequency, small amplitude f l ~ c ~ a t i o n s in the recorded
measurements before the liquid was injected due possibly to extemal electromagnetic noise
picked up by the data acquisition boards. These fluctuations were found in al1 nins perfomied
both on the ground and aboard the KC-135. It was found using Fast Fourier Transform (FFî)
method that the dominant frequency of these fluctuations was 40 Hz. The power spectra of
the fluctuations in heat flux and wall superheat data under no flow conditions aboard the KC-
135 are shown in Figures 4.3 and 4.4, respectively. In order to smooth the raw data, a three-
point moving average filter was applied in the andysis of the transient heat flux and surface
temperature characteristics. After the liquid was injected into the test section, the surface heat
flux and temperature fluctuations in amplitude increased under different boiling conditions
(from A to E).
As soon as the liquid was injected into the test section, a high surface temperature
caused dispersed fiow film boiling of liquid over the sensor surface (A - B) in Figures 4.1
and 4.2. This could be seen clearly from the top window and observed by a micro-video
camera system aboard the KC-135 and a high-speed video carnera on the ground. In that
penod, surface temperature first had a smdl sudden drop and then decreased gradually with
time. The instantaneous heat flux data showed some relatively large amplitude fluctuations in
between small-scale fluctuations. Large heat flux peaks occumng in a dispersed flow
indicate that some liquid droplets approach the hot surface but quickly bounce away or
slide over the heat flux sensor surface. Because the arnount of time for vaporizing enough
1 Run #MG-D23
Figure 4.1. A typical transient heat flux history during rewetting of R113.
O 5 IO 15 20 25 30 35
Time (second)
Figure 4.2. A typical transient surface superheat history during rewetting of R 1 13.
Run #MG-D23 (RI 13)
Figure 4.3. Typical power spectmm density of heat flux before R 1 13 injection.
14000 Run #MG-023
Figure 4.4. Typical power spectrum density of wall superheat before R113 injection.
liquid to form a vapor film at such high surface temperatures is sufficiently small, the liquid
or droplets would remain separated from the hot surface by the vapor film. Chandra and
Avedisian (1991) photographically showed that no evidence of liquid-solid contact was
found in their experiments of n-heptane droplets colliding with a hot staidess steel surface
when the surface temperature is higher than the Leidenfrost temperature. The magninide of
heat flux fluctuations in the film boiling period (A - B) is quantitatively very small compared
to those in transition boiling (B - C) and nucleate boiling regimes (C - D) in Figures 4.1 and
4.2-
When the quench front was far from the sensor location, the surface temperature
decreased gradually until the quench front came close to the sensor (A - B in Figure 4.2).
The heat flux data were characteristic of the motion of the interfacial waves which resulted in
periodical thinning and thickening of the vapor film. Adham-Khodaparast (1996)
experimentally and theoretically analyzed the liquid-vapor interfacial wave charactenstics.
He concluded that the frequency of interfacial waves increased with flow rate, but was
almost the same for different gravity levels and liquid subcooling at the inlet. The average
wave energy stored in the interface increased with gravity which together with the heat flux
data, showed the existence of a thicker and calmer vapor film in reduced gravity conditions.
Cross correlation of heat flux fluctuations showed a decrease in wavelength with flow rate, in
agreement with the theoretical effect of combined Rayleigh-Taylor and Kelvin-Helmholtz
instabilities.
There was an apparent decrease in heat flux before the quench front reached the
sensor (before point B in Figure 4.1). This decrease in heat flux could be caused by an
inverted annular flow pattern that may have existed downstrearn of the quench front. Kawaji
et al. (1985) summarized the previous visualization experiments on vertical reflooding of
tubes with water and indicated that the inverted annular 80w would occur if the liquid is
injected rapidly and remains subcooled at the quench Front. Kawaji et al. (1991) also
observed the inverted annular flow in quenching of a tube under reduced gravity. For this
fiow pattern, much more vapor with smdl axial momentum is produced at the quench front,
which can continuously thrust the liquid-vapor interface away from the hot surface,
increasing the vapor film thickness and decreasing the füm boiling heat transfer just
downstream of the quench front.
This temporary decrease in heat flux resulted in the surface temperature becoming
slightly higher than the temperature in the film boiling region ( before point B in Figure 4.2).
Adham-Khodaparast (1996) used a two-dimensional conduction mode1 to simulate the flow
of heat in the present heat flux sensor. He found that the increase in surface temperature
would be caused by the composite structure of the sensor consisting of an alurninum oxide
layer on a metal surface with significantly different thermal conductivities. This composite
structure can give rise to a different thermal response compared to that of a single component
solid, when a quench front with a steep axial temperature gradient approaches the sensor
location. Aithough the absolute values of the surface temperature measurements near the
quench front obtained with the present sensor may be slightly different from those of a
single-component solid wall, they are still valid for the study of liquid-solid contact
mechanisms and the cornparison between quenching experiments under different conditions.
With the surface temperature just upstream of the quench front decreasing below the
rewetting temperature, the rate of vapor produced and the vapor film thickness at the quench
front would be reduced. The Rayleigh-Taylor and Kelvin-Helmholtz instabilities of the
interface between the liquid and vapor streams would then readily cause a collapse of the
vapor film between the liquid and hot surface, which is considered to be the onset of
rewetting. With the collapse of the vapor film, the surface temperature profile exhibited a
knee in the slope. For some runs it was sharp, but for some other runs smoother. The
sharpness of the knee was dependent on the speed of the vapor film collapse and the
establishment of liquid-solid contacts at the quench front. If the duration of vapor covering
the heat flux sensor lasted longer than that of liquid rewetting, the knee would be smoother.
Othenvise it would be sharper.
After the onset of rewetting, the surface temperature decreased much more rapidly
through the transition boiling and nucleate boiling regimes. In heat flux signals. largescale
fluctuations appeared from the onset of rewetting through the maximum heat flux, to the
stage involving low wall superheat, conventionally regarded as nucleate boiling. When the
wdl superheat reached a sufficiently low value, the surface temperature decreased very
slowly and heat flux became almost constant, which indicates that boiling has ceased and
heat transfer from the sensor is by convection to a single-phase flow of liquid.
The heat flux fluctuations dunng rewetting at high wall superheat were an order of
magnitude larger in amplitude than those of film boiling and can only be attx-ibuted to
unstable liquid-solid contacts. As shown in Figure 4.5. the contacts were long lasting and
clearly indicated by distinct heat flux peaks, which lasted for measurable durations of time
and followed by distinct temperature dips. ïhese contacts cool the sensor surface more
rapidly so that more liquid can corne into contact with larger solid areas. The heat flux,
therefore. increases significantly while the surface temperature decreases, and vice versa,
which is a unique characteristic of heat transfer during rewetting. It is noted that the surface
temperature and heat flux fluctuated synchronously despite a difference of 3 mm in the
positions of the surface temperature and heat flux sensors. This reveals that the wetldry area
could be occupying at least an area 6mm diameter. The detailed discussion of the liquid-solid
contacts will be presented in Chapter 6.
Time (second)
Figure 4.5. Synchronized response showed by heat flux and surface temperature.
For the quenching experiments using PF5060 as a codant aboard the DC-9 aircraft
and on the ground, typicd quenching heat flux and surface superheat data sampled at 500 Hz
in one of the reduced gravity runs are shown in Figures 4.6 and 4.7. Although higher data
sampling rate was used, the heat flux fluctuations before liquid injection were much smaller
in amplitude than those in three boiling regimes. A ten-point moving average filter was again
applied to the raw data in the anaiysis of the transient heat flux and surface temperature
charactenstics. It is clear from Figures 4.6 and 4.7 that the quenching characteristics for
PF5060 are similar to those for RI 13 shown in Figures 4.1 and 4.2. The heat flux and surface
temperature fluctuations also showed synchronous variations as shown in Figure 4.8.
However, there are many differences in the behavior between these two fluids in the
quenching process, which will be discussed in the following chapters.
. . . . i l i Run #MG613
O 2 4 6 8 10 12 14 16
Time (second)
Figure 4.6. A typical transient heat flux history during rewetting of PF5060.
O 2 4 6 8 10 12 14 16
Time (second)
Figure 4.7. A typical transient surface superheat history during rewetting of PF5060.
Time (second)
Figure 4.8. Synchronized response showed by heat flux and surface temperature.
Direct measurements of both the local surface heat flux and temperame
simultaneously during rewetting are, to the author's knowledge, the fint ever in several
decades of boiling heat transfer research. Previous measurement methods using
thermocouples embedded in the plate to obtain heat flux data by inverse conduction
calculation could not detect such rapid and large fluctuations in heat flux as the present heat
flux microsensor. The data clearly show the advantage of using heat flux microsensors in
boiling heat transfer studies to elucidate the physical phenornena
4.2. B o h g Cuwes during Quenching
In the present study, the boiling curves during quenching were obtained by time
averaging heat flux and surface temperanire data over an appropriate time interval since the
experimental data and empirical correlations in the literature generally are based on time
average measurements. The effects of different averaging time on the boiling curves becarne
less as the averaging time interval was increased. Figure 4.9 shows that the heat flux for
RI 13 becarne insensitive to the number of points used in moving average if the number is
greater than 50 (0.5-second interval). A 100-point (one second interval) moving average for
R113 tests and 500-point (one second interval) moving average for PF5060 tests were
applied to heat flux and surface temperature data, which made quenching curves smoother
and the analysis easier.
It has been rnentioned in Chapter 2 that two heat flux gauges on the microsensor were
7 mm apart in the flow direction and the resistance temperature sensor was located in the
middle. If the quenching speed is high but the data sampling rate is low, as in the
experiments involving R113, this spatial difference would have little effect on the
conespondence between heat flux and temperature data in time sequence. However, this
effect was very significant in the experiments involving PF5060 because of slow quenching
speed and high data sampling rate. For instance, the heat flux and surface temperature data
shown in Figures 4.6 and 4.7 were measured with a heat flux sensor located 4rnm upstream
of the temperature sensor. Since the quench speed for PF5060 was much slower than that for
R113, as will be discussed in the next chapter, the rewetting time detected by the surface
temperature was delayed by about 3 seconds compared to that detected by the heat flux
sensor. This delay has been accounted for in constnicting the boiling curve by shifting the
temperattue data backward in time. The delay time was determined by the distance between
the heat flux and surface temperature sensors divided by the quench velocity. The
determination of the quench velocity will be presented in the next chapter.
O 5 10 15 20 25 30 35
Time (second)
Figure 4.9. The effect of time average on boiling curve for R 1 13.
The parametric effects of liquid flow rate, inlet subcooling and gravity level on the
boiling curves are shown in Figures 4.10 and 4.1 1 for RI13 and in Figures 4.12 and 4.13 for
PF5060. For both fluids, the boiling curves show similarity in the general shape between the
normal and reduced gravity conditions. Except for the lower wall superheat region, heat flux
generally increases with increasing flow rate, inlet subcooling and gravity for RI13 under
reduced gravity and some cases for PF5060 in normal gravity conditions, resulting in shifting
of the boiling curve to higher heat flux and higher wdl superheats. The reason for this shift is
Wrely that for low flow rates, low subcooling and low gravity level, the vapor layer thickness
increases in film boiling and the dnving force for the liquid to rewet the dried surface is
reduced in transition boiling and nucleate boiling at high wall superheats. Since boiling heat
Figure 4.10. Boiling curves measured in 1-g for RI 13.
Figure 4.1 1. Cornparison of boiling c w e s measured in p-g and 1-g for R113.
Figure 4.12. Boiling curves in 1-g with low and hi& inlet subcooling for PF5060.
Figure 4.13. Boiling curves in p.-g and 1 -g with high inlet subcooling for PF5060.
transfer is more hydrodynamically controlled at sufficiently high wall superhea~, the effects
of reduced liquid momentum and gravity, and an increase in vapor generation rate are
amplified at high wall superheats. Kawaji et al. (1991) and Westbye et al. (1995) also showed
the sarne trend in rewetting of a hot q u m tube and a metal tube with R113, respectively.
High inlet subcooling (ATrob = 33 OC) combined with a high flow rate (A400 k&s)
has a strong effect on boiling curves in depressing heat flux in nucleate boiling as shown in
Figures 4.12 and 4.13. The boiling curves for PF5060 with inlet subcooling of 18 OC for ail
flow rates at 1-g are similar to those for R113 with inlet subcooling of 25 O C in nucleate
boiling at lower wall superheats. The heat flux decreased gradually rather than sharply with
decreasing wall superheat. However, the heat flux in nucleate boiling regime for PF5060
with high inlet flow rates (>1400 k&s) and hi@ inlet subcooling (ATrub = 33 OC) in 1-g
and p.-g conditions decreased sharply with decreasing wall superheat. The heat flux in
nucleate boiling was strongly depressed under those flow conditions.
Comparing the boiling curves for R113 and PF5060, it is found that the boiling
curves for R113 shifted to higher wall superheats compared to those of PF5060. This shift is
believed to be caused by the therrnophysical properties of R113, mainly the density ratio of
liquid to vapor and latent heat. The lower density ratio and higher latent heat of R113
compared to PF5060 made it easier for the liquid to rewet the hot surface even at higher wall
superheats, which will be discussed in the following chapter(s).
Some researchers such as Yilmaz and Westwater (1980), and Peng et ai. (1992) have
found that high liquid subcooling and flow velocity can dramatically alter the traditional pool
boiling curve. Transition boiling regirne tends to disappear with increasing flow rate and
subcooling, and furthemore, the minimum heat flux will reach the critical heat flux. This is
consistent with the effects of Aow rate, inlet subcooling and gravity observed in the present
experirnents.
It was found from present and previous experiments that there is a slight effect of
liquid subcooling, fiow rate and gravity on low wall superheat nucleate boiling. However,
heat flux at higher wall superheats in transition boiling significantly increases with increasing
liquid subcooling, flow rate and gravity. The reason for decreasing wall heat flux with
increasing wall superheat is the accumulation of coalescing bubbles which hover over the
boiling surface. This vapor cluster above the boiling surface can interfere with the further
release of bubbles from the boiling surface, which increases the thermal resistance for boiling
heat transfer. Increasing flow rate, subcooling and gravity c m decrease this effect. If the
liquid could quickly refill the space from where the bubbles have departed, heat flux would
increase with increasing wall temperature. If this happens, the boiling curves could behave
like those shown in Figure 4.14.
Figure 4.14. The discussion of boiling mechanism.
At low wall superheats (O - A), the active nucleation sites increase in nurnber density
with increasing wall superheat. There is no interaction between the bubble releasing colurnns.
When the wall superheat rises to a certain value (point A), active nucleate sites increase so
that neighboring bubble columns can coalesce or merge at certain places, where larger bubble
columns can be seen. As bubbles depart from the nucleation sites, new bubbles grow at the
sarne time, and the waiting time approaches zero. With a further increase in wall superheat
(A - B), bubbles will accumulate above the boiling surface and form a vapor cluster due to
the effect of the wake of a leading bubble on the trailing bubble. The formation of a vapor
layer leads to local dryout and reduction in the boiling heat transfer rate on the surface so that
the slope of the heat flux curve decreases with a M e r increase in the wall superheat until
the maximum heat flux (B) is reached. In this region, a vapor cluster departs and another
forms immediatel y, resulting in intermittent local dry out and rewe tting.
This indicates that boiling heat transfer has changed from therrnodynamically
controlled boiling on the surface to hydrodynamically controlled heat transfer above the
surface. It may be useful to consider that the ability to generate vapor bubbles (AGV) is
competing with the ability to remove the vapor (ARV) from the surface at higher wall
superheats. Both increase with increasing wail superheat. From O to A, ARV is much greater
than AGV. But AGV is approaching ARV from A to B and equals ARV at point B. In the
region fiom A to B, intermittent rewetting of local dryout areas cm be maintained. However,
when the wall superheat is higher than at point B, AGV is greater than ARV so that the
dryout area spreads to a larger size than in the A - B region, which causes the heat flux to
cirop on the average (instantaneous maximum heat flux can be larger than in the region, A - BI.
For heat flux controlled heating systems, the dryout area could spread to the whole
heating area and a vapor layer forrns above the heating surface. For temperature controlled
systems, it is found from previous pool boiling experiments (Kalinin et al. 1987) that the
frequency of vapor cluster release or liquid rewetting-dryout area increases with increasing
wall temperature until point C and then decreases until point D. The reason could be that the
liquid-solid contact frequency is wall temperature controlled from point B to point C and
Iiquid rewetting ability controlled from point C to point D. The lower wall temperature in the
range from point B to point C causes longer waiting period to produce a dry area than that in
the range from point C to point D. So, the wail temperature is a dominant factor for boiling in
this range because liquid has a strong ability to rewet the dried area. With increasing wall
temperature from B to C, the waiting period decreases so that liquid-solid contact frequency
increases. However, the situation is reversed in the range from C to D. The ability of liquid to
rewet the dried area becomes a dominant factor for boiling because a large arnount of vapor
is generated and dried area can be produced easily in such a high wail temperature range.
With the increasing wdl temperature from C to D, the dried area becomes larger and the
ability of the liquid to rewet the dried area is reduced so that the liquid-solid contact
frequency decreases.
More detailed discussions on particular topics such as rewetting temperature, quench
velocity, maximum heat flux and three boiling modes will be presented in the following
c hap tea.
4.3. Flow Boiling Visualization
43.1. Quenching experiments with heat flux indicator
After the heating surface was heated to about 200 O C , the power was shut off.
Subcooled liquid R113 was circulated and injected into the test section to quench the hot
surface. The high speed video camera with a 186 Hz frame rate and 11500 shutter speed was
used to record the boiling process and heat flux level at the same time. Because of the limited
memory of the high speed video camera system, the boiling process was recorded for only 3
seconds.
Figure 4.15 shows an image in the film boiling regime. Since a very thin vapor film
was formed between the liquid and solid surface, the heat flux sensor could be seen from the
top and heat flux was between 100 and 200 kw1mz as indicated by the heat flux indicator.
While the liquid began rewetting the hot surface, the heat flux sensor still could be viewed in
transition boiling and the heat flux increased as shown in Figure 4.16, reaching between 300
and 400 k ~ l m ' . When the heat flux level exceeded 1ûûû k ~ l m ' , the heat flux sensor was
hlly covered by the vapor (Figure 4.17). After the wall temperature dropped below the
temperature corresponding to the maximum heat flux, bubbles could be observed on the
sensor and the heat flux was reduced to about 700 kw1m2 (shown in Figure 4.18). Since there
was too much vapor generated during the quenching process, it was difficult to view the
entire boiling process in detail.
43.2. Experiments on quenching of a quartz tube with water
For quenching of a quartz tube with distilled water at a high flow rate, Figure 4.19
shows a whole view of the quench process in 13.44 ms intervals at a frarne rate of 744 Hz
and shutter speed of 1110,000 sec". Because of high flow rate and wall temperature, the
liquid downstream of the quench front formed dispersed film boiling. The quench front
Figure 4.15. The quenching of R- 1 13 image at film boiling regime ( 1 -g).
Figure 4.16. The quenching of R- 1 13 image at transition boiling regime ( 1 -g).
Figure 4.17. The quenching of R- 1 13 image at maximum heat flux ( 1 -g).
Figure 4.18. The quenching of R- 1 13 image at nucleate boiling regirne ( 1 -g).
in the middle moved slightly faster than that at the side. Near the quench front, the vapor
film locally broke d o m over a smaii area (spot 1 in Figure 4.19 (a)) and was rewetted
(Figure 4.19 (b)) by the liquid. As the quench front further advanced, the rewetted area grew
in size in the rniddle area of the wdI (spot 1 in Figure 4.19 (c) and (d)). Then the liquid
rewetting the dry area was evaporated and vapor hovered over the area again as show in
Figure 4.19 (e). With further propagation of the quench front, the boiling mode at spot 1
changed from transition boiling to nucleate boiling (Figure 4.19 (f) to 0)). Because of low
themal mass of the thin-wailed quartz tube, only one cycle (5.4 ms) of liquid-solid contact
was required to quench the surface. If the thermal rnass of the hot surface were much greater,
then the number of rewet-dryout cycles required to quench the surface would be much
greater than one.
(cl
ti)
Figure 4.19. The vapor film broke behind the quench front (1-g).
Film Boiling, Rewetting Temperature and Quench Velocity
5.1. Film Boiling
Film boiling is usually distinguished by low heat transfer coefficients and high
surface temperatures compared to nucleate boiling. The heat transfer coefficient, hlb, in the
film boiling regime is particularly important in determinhg the total quench time of a hot
surface. Due to the presence of liquid-vapor interfacial waves, the value of hla for R113
fluctuated with the surface temperature slightly decreasing until the quench front reached the
heat flux sensor as shown in Figure 5.1. The effects of flow rate, liquid subcooling and
gravity on film boiling heat transfer for R113 c m be seen from the time-averaged values
shown in Table 5.1. The value of hfi increases with increasing flow rate, liquid subcooling
and gravity, which is expected and has been observed during quenching of a tube by Westbye
et al. (1995) and other researchen. However, Run MG-D25 shows no change in heat transfer
coefficient. This can not be explained at the present time due to a lack of sufficient reduced
gravity data. For the experiments using PF5060 as a coolant, a reliable set of film boiling
heat transfer data could not be obtained because the heat flux microsensor fabricated on a
stainless steel disk showed large offset in film boiling regime due to an unknown reason.
Table 5.1. Film boiling heat transfer coefficients in p-g and 1-g experiments. (R113)
MG-D23 MG-D25 MG-D20 NG-G2 NG-GS NG-G 10 NG-G6 NG-G7
0.8 (GS) O. 194 0.243
0.7 (G IO) 0.294 O. 1 19 0.194 0.289 0.222 0.298
In Westbye et al.'s (1995) work, the values of hfi in p.-g were found to be much Iess
than those obtained in 1-g and the ratio of hP in p.-g to 1-g ranged from 0.15 to 0.60. On the
ground, because gravity causes flow stratification in the horizontal test section, and reduces
the vapor film thickness above the hot surface, higher heat transfer coefficients are obtained
in the film boiling regime. In reduced gravity, however, this effect is absent and the vapor
film becomes thicker. In the present results, the ratios of hfi in p.-g to 1-g for two pairs of
runs with similar inlet flow rate and liquid subcooling are 0.7 and 0.8, which are higher than
Westbye et al.'s (1995) results. Probably this is caused by higher flow rates in the present
experiments and a different heat flux measurement method, i.e., inverse conduction method
applied to tube temperature measurements, used in their experiments.
Run MG-023
6 8 10 12 14 16
Time (second)
Figure 5.1. Typical film boiling heat transfer coefficient profile.
Many theoretical and experimental approaches have been used to study film boiling
heat transfer in the pst several decades, and most of the studies focused on pool boiling, as
reviewed by Carey (1992). Adham-Khodaparast (1996) surnmarized the previous work on
film boiling heat transfer mechanisms and pointed out that the analysis of film boiling heat
transfer cm be grouped into two major groups, the laminar and turbulent vapor film flow
approaches. The 1amina.r film mode1 is valid for film boiling on short heaùng surfaces or over
the short distance from the Ieading edge. For longer surfaces, the prediction would result in
underestimation because the vapor flow changes from laminar to turbulent. In the present
study, the heat flux microsensor was placed 50 mm downstrearn of the nlet and the flow of
vapor is believed to remain in the laminar flow regime. Therefore, the previous correlations
based on a lamuiar film model could be applied in the present study.
Many correlations have been developed for flow film boiling heat transfer from
experimental data obtained with different fluids. The Martinelli parameter and superposition
principle are commonly used in correlating the experimental data In the present study,
Bromley's (1953) correlation developed from a laminar film model was compared with the
present data in 1 -g conditions.
Although Brode y's ( 1953) correlation was developed for upward forced convection
film boiling across a horizontal tube, it is assumed to be valid for the present case because of
very thin vapor film thickness. Bromley's correlation for flow film boiling is as follows:
for
and for
w here h '1" = hiv 114.4 cp.. ( T w - Tm )/ hhl2
Since film boiling in the present study occurred close to the minimum film boiling
point and hydrodynamic instability c m be considered to be the main factor causing vapor
film collapse, the critical or most dangerous wavelength of a disturbance in the liquid-vapor
interface c m be chosen as the characteristic length L, (Carey, 1992)
where o is surface tension (N/m).
The average heat transfer coefficients in 1-g caiculated by equation (5.1) using the
value of inlet flow velocity for the run NG-G2 and equation (5.2) for other runs are Iisted in
Table 5.1. The predictions are very close to the present experimental results in 1-g conditions
with differences ranging €rom 3% to 1 1 %.
The heat fluxes calculated by equations (5.1) to (5.4) are shown in Figure 5.2 and
compared with the results from two pairs of nuis under both gravity levels. Due to smaller
wall superheat and large fluctuations in heat flux, it is hard to say how well the predictions fit
the experimentd data in 1-g conditions. However, the predictions are very close to the mean
values of the experimental data.
It is found from Figure 5.2 that the heat fluxes for p-g experirnents are at lower wall
superheat ranges than those for 1-g experiments. The question would be whether the lower
film boiling heat transfer coefficients in p-g discussed above are caused by lower wall
superheat. The cornparison of hk in p-g with that in 1-g predicted by Bromley's correlations
presented in Table 5.1 clearly shows lower hk is obtained in p-g. However, a definitive
conclusion can not be derived from the RI13 results due to a lirnited amount of reduced
gravity data available.
5.2. Rewetting Temperature
The local wail temperature at the onset of rewetting is very important for theoretical
modeling and engineering applications. Many definitions have been used in the literature,
such as rewetting temperature Tm, apparent quenching or rewetting temperature Ta,
minimum film boiling temperature T-, Leidenfrost temperature, etc. This can be confusing
and do not always represent the sarne physical phenornenon. Many researchers used one of
the above definitions according to their expenmental configurations.
Barnea and EIias ( 1994) have presented the following argument about the first two
terms: the apparent rewetting temperature, Tw, is defined as the intersection between the
tangent line to the temperature-time curve at the point where its slope is the largest, with
the tangent to the curve before quenching. T, marks the onset of rapid surface cooling
caused by an enhanced rate of heat transfer that does not necessitate liquid-solid contact. The
rewetting temperature ,Tm, on the other hand , is the temperature at which a triple interface of
vapor-liquid-solid is formed. This temperature is difficult to define from the measured
temperature - time curve, and in their study, they considered it as the highest temperature at
Figure 5.2 (a). Comparison of film boiling heat transfer data for RI13 with the results ~redicted bv eauations (5.1) to (5.41. U;, = 0.38 m/s.
1 20 1 40 160 180 200 220 240 260
T, - T,, (OC,
Figure 5.2 (b). Comparison of film boiling heat transfer data for R113 with the results predicted by equations (5.1) to (5.4), Uin = 0.88 &S.
100 -
80 - f 3 60 - Y Y
NG-G 1 0 (0.93 mis)
MG-020 (0.84 mis)
m 40
20 -
which the slope of the surface temperature vs. time curve first exceeds an arbitrary value of
500 OCls. This is the traditional way to detennine the rewetting temperature from measured
temperature-time curves. Hung et al. (1994) identified the minimum boiling temperature,
T,, with the Leidenfront temperature which is defined as the maximum possible liquid-solid
contact temperature, because the inverted annular flow generdly occurred at the minimum
heat flux conditions in their experirnents. houe and Tanaka (1991) defined the apparent
rewetting temperature, T,, according to different cases of boiling curve configurations. In the
case of dispersed flow film boiling, Ta,, is the temperature at which the boiling curve begins
to depart from the trend of film boiling; and in the inverted annular flow, it is the temperature
at which the heat flux increases sharply; and in other cases, it is the temperature at the
minimum heat flux point. They distinguished the apparent rewetting temperature, Ta, from
the rewetting temperature, Tm, in the inverted annular flow region, and defined it as the
temperahue at which the heat flux begins to increase away from the trend in the hi@ wall
superheat region.
In general al1 of them used thermocouples embedded inside the heated wall to
estimate the surface temperature and to calculate the heat flux by inverse conduction
routines. Actually, they could not obtain the actual surface temperature at the onset of
rewetting and detect the abrupt change in heat flux as liquid-wall contact was initiated. In the
present study, the initiation of film boiling collapse and the onset of rewetting have been
detected by direct rneasurements of heat flux and surface temperature using the micro heat
flux sensor and RTS, as can be seen in Figure 4.1 where large fluctuations started to occur in
heat flux and surface temperature.
It can also be seen in Figure 4.1 that heat conduction in the direction paralle1 to the
surface can apparently cause instantaneous revend in heat flux at the onset of rewetting as
the quench front arrives at the heat flux sensor. Two dimensional heat conduction in the heat
flux microsensor has been studied by Adharn-Khodaparast (1996), as mentioned in Chapter
2. Also many researchers, such as Chen, Lee et al. (1979), Cheng et al. (1978)- Huang et al.
( 1993, 1994), and others have noted and studied two dimensional heat conduction in the solid
during quenching.
In the present study, the temperature at the onset of rewetting is referred to as the
rewetting temperature, T,, which is determined just when the large heat flux fluctuations
first appear. Since the heat flux sensor and surface temperature sensor are not located exactly
at the same location, rewetting would occur on the senson at different times. This time lag
has k e n accounted for by considering the quenching velocity, and calculated by dividing the
distance between two sensors by the quench velocity. in determining the rewetting
temperature, Tw. It is noted here that the quench velocity will be discussed fully in the next
section.
Since there are many parametnc effects on rewetting temperature, such as initial wall
superheat, inlet liquid subcooling, flow rate, gravity level, and pressure, we will study only
some of them. In the present work, the experiments were conducted at near atmospheric
pressures, so the effect of pressure cm not be addressed. However. many researchers have
previously reported a positive effect of pressure on rewetting (Huang et al. (1 994)).
Figure 5.3 shows the effect of the initial wall superheat (Tw - T,J on the rewetting
temperature for RI13 with subcooled inlet and in normal gravity conditions. Most of the data
are scattered in a curved band, except for two data points. lmmediate rewetting occurred for
the first case where Tw is well above the band. For another case, showing T, well below the
band, film boiling lasted much longer prior to rewetting. which indicates greater stability of
the vapor film. It should be noticed that these two special cases occurred at medium fiow
rates in the present experiments. Chen, et al. (1979) in a study of rewetting of a hot circular
pipe with water at atrnospheric pressure, observed that once the initial wall temperature
exceeds 650 OC (wail superheat of 550 OC), the rewetting temperature rernains almost
constant. This phenomenon did not occur under the present experimental conditions, since
the wall superheat ranged from 166 to 240 OC. However, there is a slight trend to
asymptotically approach constant Tm, as the initial wall temperature increases, as can be seen
in Figure 5.3.
Figure 5.4 shows the rewetting wall superheat, AT,, or (Tw - T,J for R113 in
normal gravity conditions with subcooled and saturated liquid at the inlet. The rewetting wdl
superheat, AT,, for subcooled liquid is higher than that for saturated liquid at the inlet. With
increasing flow rate, AT, increases for the subcooled liquid but only slightly for saturated
liquid. Ueda et al. (1983) performed rewetting expenments with an upward flow of
subcooled R113 through a cylindrical tube equipped with a heated copper block at a system
pressure of 0.32 MPa They showed that signifiant changes occur in the rewetting
temperature with the inlet quality or subcooling. From their data, it could also be seen that
the rewetting temperature increaîed slightly with increasing flow rate for the d e t quality
between -0.1 and 0.15, which is in accordance with the present results for saturated inlet
case.
I . A ~ T ~ = 1 6 8 " ~ :-- A a..' ~T,,,=178"~
A A T ~ = I ~ O ~ C
..S. ATw=201 OC
ATw over 206 OC
Figure 5.3. The effect of initial wall superheat on rewetting temperature for RI 13.
1 -g, AT, = 25 O C A
O
O 0
Figure 5.4. The rewetting superheat for R113 in p-g and 1-g.
On the other hand, houe and Tanaka (1991) conducted experiments similar to Ueda
et aL's (1983) using a stainless steel test section and hot patches at the inlet and outlet of the
test section, and they found that quality and flow rate only slightly affected the rewetting
temperature over al1 the test ranges.
Contrary results were also obtained for water in rewetting studies made by Iloeje et.
al. (1975), Groeneveld and Stewart (1982) and Cheng et al. (1985). Hoeje et al. (1975)
reported significant effects of inlet quality and water flow rate on the rewening temperature
in a circular tube at a pressure of 6.89 MPa. Groeneveld and Stewart (1982) found that both
the inlet water flow rate and quality did not have any significant effect on the minimum film
boiling temperature in the saturated region, but a large increase in the minimum film boiling
temperature was observed for increasing liquid subcooling in the subcooled region. However,
Cheng et al. (1985) performed experiments sirnilar to Groeneveld and Stewart's and showed
that the minimum film boiling temperature increases with increasing mass velocity and
decreasing quality. So far, it has been very hard to draw definitive conclusions on the effects
of flow rate and subcooling andor quality on rewetting temperature or minimum film boiling
temperature due to a limited arnount of expenrnentai results available.
Next, Westbye et al. (1995) reported a significant effect of gravity on the rewetting
temperature and a slight effect of flow rate at both gravity levels in rewetting of a hot,
horizontal thin-walled stainless steel tube at amospheric pressure using R- 1 13. However, the
effect of 80w rate on rewetting superheat in 1-g with subcooled inlet in his study was
different from the present results, as presented above. The reason is that the inlet velocity
range (from O. 1 to 0.5 m l s ) in his study was only half of the present range (from O. 1 to 1 .O
d s ) . Therefore, their data could not show the significant effect of flow rate on the rewetting
superheat as shown in Figure 5.4. The rewetîing superheat for R113 in p-g with 25 OC
subcooling is about 10 OC higher than that in I-g with saturated inlet and about 10 to 45 OC
lower than that in 1-g with the same inlet subcooling.
In inverted annular film boiiing flow, the interface between the liquid and vapor is in
constant motion and a Kelvin-Helmholtz type instability will set in if the two phases have a
large enough relative velocity. The amplitude of interfacial waves will grow until the liquid
contacts the solid surface. If the initial surface temperature is higher, more vapor will be
generated which will enhance heat transfer at the interface and increase the vapor velocity.
Therefore, it wiil cause greater interfacial instability and breakdown of the vapor film at a
higher wall superheat. At higher liquid flow rates, the vapor film thickness would decrease
which leads to rewetting at a higher wall temperature. The vapor film thickness increases
with decreasing inlet subcooling and gravity level, which leads to more stable film boiling
and a reduction in the rewetting temperature.
The rewetting superheats for PF5060 at both gravity levels with subcooled inlet
conditions are shown in Figure 5.5. In reduced gravity, the rewetting superheats for the
higher initial wall superheat were significantly higher than those for the lower initial wdl
superheat, and close to those obtained in normal gravity conditions. The rewetting superheat
data in 1-g showed a srnall effect of inlet liquid subcooling. In contrast with the results for
R113, the rewetting superheat for PF5060 showed no effect of flow rate over the flow
velocity ranging from 0.4 to 1. I mls in the present study. The reason could be due to the
different thermophysical properties. It is noticed that the latent heat of PF5060 (hi, = 84.6
I d k g ) is smaller than that of R i 13 (hl" = 144.1 Id/kg), which would cause more vapor to be
generated and therefore a thicker vapor film to be formed at the quench front for PF5060.
Figure 5.5. The rewetting superheat for PF5060 in p-g and 1-g.
A comparison of the rewetting superheat between the two fluids is shown in Figure
5.6. The rewetting superheat of RI 13, ranging from 140 O C to 200 OC, is about twice that of
PF5060, ranging from 60 OC to 100 O C . Although PF5060 has smaller latent heat, its vapor
density (1 1.2 kg/m3) is much greater than that of RI13 (7.3 kg/m3). Physically, a liquid
having a smailer latent heat will be evaporated more per unit heat input, so that a larger
amount of vapor will be generated near the hot surface. Also, the larger the vapor density is,
the larger the vapor momentum would be to push the liquid away from the hot surface. Both
of these factors contribute to delaying of the vapor film collapse for PF5060 until Iower wall
superheats are reached compared to PF5060.
Figure 5.6. Cornparison of rewetting temperatures for R113 and PF5060.
Although it is quite complicated and difficult to Fully analyze the rewetting
phenomena, many researchers have tried to predict the minimum film boiling temperature
and rewetting temperature for flow boiling and quenching by theoretical models and
experimenial correlations. Berenson (196 1) extended Zuber's vapor escape mode1 to analyze
the minimum heat flux condition in steady film boiling over a flat horizontal surface. He
described the heat transfer through the vapor film as a pure heat conduction problem. The
vapor film thickness was also calculated according to the hydrodynamic stability condition
and by incorporating some empirical interpolation. Berenson (1961) obtained the following
correlation to predict the minimum film boiling temperature for pool boiling:
The minimum waii superheat calculated from equation (5.5) for R 1 13 at atmospheric
pressure is 67.8 O C , which is much lower than the present data and is noted here as the
Berenson's minimum wall superheat, ATmBep Lienhard (1976) applied the Maxwell-Van der
Waals theory to correlate the available data and recommended a simple correlation for the
maximum liquid-solid contact temperature in pool boiling as follows:
AT, = 0.905 - Tm + 0.095 T> (5.6)
where, Tm = T, / Tc, T, is saturation temperature and Tc is the critical temperature. The
lirniting liquid superheat calculated by (5.6) for RI13 at atmospheric pressure is 123 OC,
which is still lower compared to the present data.
noeje et al. (1975) conducted vertical flow boiling expenments with water in an
inconel tube and observed minimum film boiling superheats asymptotically approaching
certain values. They intuitively expected that this asymptote would be close or equal to a
pool boiling value. From this expectation, they correlated their data in the following
empirical fom:
where X is quality, G is mass flux, A, B, m and n are constants.
As mentioned above and also seen from Figure 5.4, the flow rate effect on the
rewetting superheat for R 1 13 varies with different subcooling and gravity levels. When the
vapor film thickness increases for a saturated inlet condition under normal gravity and
subcooled inlet condition under reduced gravity, the vapor velocity and the shear at the
interface both decrease. Therefore, the disturbances caused by the increased liquid flow rate
have relatively little effect on Kelvin-Helmholtz instability and this leads to a smaller effect
on rewetting temperature for the two conditions. This indicates that the effect of mass
velocity is strong when the vapor film is relatively thin, which occurs only under subcooled
flow conditions in normal gravity.
In the present study, the rewetting superheats of RI 13 for subcooled inlet flow in 1-g
conditions also showed a similar trend of decreasing to certain values with decreasing idet
flow velocity. Thus, a correlation similar in form to equation (5.7) was used to best fit the
present data for RI 13 injection with 25 O C inlet liquid subcooling and the following
correlation was obtained,
AT, =AT,,,,, (liU.668~"~~)
where ATmBe, = 67.8 OC and G = mass flux (kg/m2s).
The predicted result is shown in Figure 5.4 by a solid line and good agreement with
experimental data could be achieved (R' = 52%). The exponent, 0.134, is smaller than 0.49
obtained by Iioeje et al. (1975), but is close to the value of 0.135 used in the correlation
given by Kim and Lee (1 979).
Here, cp, k, p are the specific heat, conductivity and density of the wall, respectively,
8 is the wall thickness,
z is the elevation and
G is the mass flux.
Kim and Lee's (1979) correlation is an empirical correlation derived from
dimensional analysis with the constants obtained from bottom flooding water tests at
atrnosphenc pressure. However, equation (5.9) involved wall superheat, AT* = Tw - Tm. For
a particular inlet subcooling (TsarTh) and elevation 2, the following form of the correlation
can be obtained from equation (5.9),
where C is a constant.
Figure 5.7 shows the present data in terms of the parameters, (Tm - T,JI(T, - T,J~.~~~ and mass flux, G. It is found that the ratios of rewetting superheat to wall superheat
for RI 13 and PF5060 remain constant with increasing mass flux. The value of the ratio is
about 2.7 for R113, which is higher than an average value of about 1.9, for PF5060. This
inconsistency shows the differences in behavior of R113 and PF5060 from water over the
present flow rate ranges.
Figure 5.7. The parameter, A T ~ / A T ~ VS. G for RI13 and PF5060.
5.3. Quench Velocity
As descnbed in the last chapter, a vapor film collapses when the surface temperature
decreases to a certain value so that the vapor-liquid interface loses stability. Quench velocity
is defined as the velocity at which the quench front propagates dong the hot surface. Quench
velocity can be detennined by measuring the time difference between the minimum heat
fluxes measured by the two heat flux senson, which are 7 mm apart in the flow direction.
The time difference in RUN MG-D23-207-4.5-22, for exarnple, is 0.65 second as shown in
Figure 5.8. The quench velocity is then the distance between the two heat flux sensors, 7 mm,
divided by 0.65 second, or 10.8 mm/s. Table 5.2 lists the quench velocities obtained in
microgravity and normal gravity experiments for R 1 13 for different initial wall temperature,
inlet flow rate and subcooling.
800 Run MG-D23 1
Time (second)
Figure 5.8. Determination of quench velocity.
Figure 5.9 shows that the quench velocity for R113 apparendy increases with
increasing inlet flow rate under both rnicrogravity and normal gravity. This result is
consistent with the normal gravity results reported previously by many researchers: Chan and
Bane rjee (1981), Lee and Shen (1987), Lee and Kim (1987). Westbye et al. (1995), and
Bamea and Elias (1994). The increase in inlet flow rate induces the liquid fraction
downstrearn of the quench front to increase and enhances the convective heat transfer to the
vapor. This results in faster cooling of the hot surface ahead of the quench front, that leads to
a higher quench velocity.
Table 5.2. Quench velocity for R 1 13 in p.-g and 1-g experiments.
MG-D23 207 22 364 10.8 0.0296 MG-D25 222 22 569 12.7 0.0224 MG-D2O 2 13 2 1 835 17.5 0.02 10 NG-GO 220 20 275 35.0 O. 127 NG-G 1 222 22 333 35.0 0.105 NG-G2 237 2 1 208 5.8 0.0280 NG-G3 235 23 125 5.8 0.0466 NG-G4 247 2 1 125 4.7 0.0374 NG-GS 248 22 397 10.0 0.0 122 NG-G9 253 2 1 9 17 17.5 0.0 189 NG-G 10 288 25 9 17 5.8 0.0064 NG-G8 257 31 134 5.7 0.0425 NG-G6 248 3 3 52 1 11.7 0.0224 NG-G7 249 32 938 23.3 0.0249 NG-GS2 200 44 332 17.5 0.0527 NG-GS6 200 45 468 17.5 0.0374 NG-GS7 202 46 679 23.3 0.0344 NG-GS8 200 47 945 1 16.7 0.123
Figure 5.9. The quench velocity for R113 in p-g and 1 -g.
25
: 2o :
15 -
RI13 A p-9, ATw= 25 O C
1-g, AT,,, 25 O C
1-g.ATw=150C A
10 -
5 -
A A rn
? =
O -- v I
Many investigators have found that the quench velocity, IIq, decreases with
increasing initial wail temperature, T., and decreasing liquid inlet subcooling. These
conclusions are consistent with Our experimental data.
For PF5060, it is shown in Figure 5.10 that flow rate, subcooling and gravity had little
effect on quench velocity when the inlet velocity was less than 0.7 d s . However, the quench
velocity increased with increasing inIet velocity and subcooling when the inlet velocity was
greater than 0.7 m/s.
In comparing the quench velocity data between R113 and PF5060, the quench
velocity for R113, ranging from 5 to 23 rnm/s, was found to be higher than that for PF5060,
ranging from 1 to 4 mm/s (shown in Figure 5.1 1). The reason is sirnilar to that given for the
rewetting superheat in the last section, however, the matends of HFMS disk, copper and
stainless steel with different thermal conductivities, could also af'fect the quench velocity as
described in Appendix I(1). Nevertheless, the above conclusion is still valid in the present
work. A detailed discussion on the efiect of the KFMS disk material on other quenching
characteristics are presented in Appendix I( 1 ).
Barnea and Elias (1994) found that the ratio of quench velocity, CIq, to inlet velocity,
LI, varied between 0.2 and 0.8 depending on the initial surface temperature and the inlet
liquid temperature, for water quenching of a vertical tube on the ground. As shown in Figure
5.12 for the present work, the ratio, U&-,,, for RI13 and PF5060 also did not Vary much for
most of the runs under both gravity levels. The ratio of quench velocity to inlet velocity for
RI13 and PF5060 is about 0.025 and 0.002, respectively for the inlet velocity between 0.2
and 1 mis. This is consistent with Barnea and Elias' results.
A cornparison between the runs for R113 perfomed in p-g and 1-g is difficult to
perfonn due to a very lirnited amount of microgravity data available and the use of a much
higher initial wall temperature in normal gravity experiments. However, comparisons of the
data in the runs GO and G l in Table 5.2 with those in reduced gravity and in normal gravity
with saturated inlet showed lower quench velocities in the absence of gravity level and liquid
subcooling. AIso, the results fiom the experiments for PF5060 show that the quench velocity
under microgravity is lower than that under normal gravity (Figure 5.10). Since the
temperature required to initiate rewetting is substantially lower in microgravity than in 1-g,
the time taken to quench the entire hot surface would be much longer in rnicrogravity.
Figure 5.10. Quench velocity for PF5060 in p-g and 1 -g.
Figure 5.1 1 . Cornparison of the quench velocity of RI 13 and PF5060.
Figure 5.12. The ratio of quench velocity to inlet velocity for RI 13 and PF5060.
Transition Boiling Heat Transfer
6.1. The Mechanism of Transition Boiling
6.1.1. Background
As mentioned in Chapter 4, transition boiling starts at the onset of rewetting and ends
as the surface is completely rewetted. Because it is between film boiling and nucleate boiling
regimes, its characteristics include the features of film and nucleate boiling heat transfer.
Transition boiling is the least understood regime in ail of the boiling repimes. The reason
is that transition boiling is regarded as technologically less important than nucleate or f h
boiling. In addition the lack of understanding is certaidy due to the complex mechanism
involved and the difficulties encounteied in experimental studies. In recent years, the interest in
this boiling regime has increased mainly in connection with the safety analysis of nuclear
reacton. In hypothetical loss of coolant accidents the transition region is traversed in a transient
process. Other quenching processes, e.g. in material processing, also go through the same
boiling curve.
Reliable prediction methods for transition boiling heat transfer are also required to
design high-performance evaporators heated by a liquid or a condensing fluid. Such heat
exchangers cm be operated in the transition boiling mode without the danger of instabilities
because the heat transfer process is temperature controlled.
6.1.2. Past modeling efforts under pool boiling conditions
The transition boiling mechanism in pool boiling is the bais for understanding the
phenornenon and its dependence on various parameten, and for developing theoretical models
and corre1ations.
The experimental results on the transition boiling mechanism and the estirnates of heat
transfer rates show that at any time, some parts of the heating surface are wetted by the liquid
and the remainder is covered by a vapor film. In this case, each point of the heating surface is
alternately in contact with the liquid and vapor. Since the rate of heat transfer to the Liquid is
higher than that to the vapor, the processes at the points of wall-liquid contact are dominant in
transition boiling.
Bankoff and Mehra (1962) proposed that transient conduction is the principal heat
transfer mechanism during liquid-solid contact. On the other hand, Kano and Yokoya (1968)
proposed that boiling heat tramfer at high heai fluxes is the dominant heat transfer mechanism
d h n g contact. At high heat flues, boiling heat transfer is characterized by the existence of a
Liquid film between the heating surface and large mushroom-like bubbles. The bubble is
nourished with vapor from many vapor stems which bridge the surface with the bubble through
the liquid film. In the nucleate boiling regime, the surface does not dry out. When critical heat
flux is reached, the liquid film evaporates away just as the bubble leaves the surface. In the
transition boiling regime, the liquid Hm evaporates away and the surface is dried out for a
p e n d of time. They assumed that the bubble period remains at that of the critical heat flux and
that the nucleate boiling c w e can be extrapolated into the transition boiling regime.
Consequently, they were able to predict the liquid füm thickness and effective heat flux at a
given surface temperature.
Kostyuk et al. ( 1986) proposed a semiempirical model for transition boiling. The model
involves transient conduction, boiling incipience and heat transfer during liquid-solid contact.
The termination of contact is caused by the coalescence of bubbles as they reach a critical
popuiation.
Fanner et al. ( 1987) developed a model for liquid-solid contact in the transition and film
boiling regimes. The mode1 is sirnilar to that of Kostyuk et al. (L986), and the contact is
modeled by incorporating transient conduction, boiling incipience and heat transfer, and
microlayer evaporation. The microlayer is the liquid film left beneath a fast growing bubble.
When the number of bubbles formed per unit area is large enough, the bubbles will coalesce,
and subsequently force the bulk liquid to retreat and leave the liquid film below the bubble.
Recently, Pan (1989) developed a model based on previous models and experimental
observations of temperature fluctuations. The liquid-solid contact process is divided into five
penods, 1) bubble departure and vapor conduction; 2) iiquid-solid contact and transient liquid
conduction; 3) boiling incipience and heat transfer; 4) macrolayer evaporation, which is the
sarne as that of Katto and Yokoya's model (1968); and 5) vapor covering and vapor conduction.
In addition to the models mentioned above, Hsu and Kim (1988) proposed a statistical
approach to treat transition boiling, in which the transition boiling cuve is simulated by a
Poisson distribution.
6.13. Past experimental work on üquid-sdid contact
It is assumed that in the transition boiling regime the heat flux cari be considered as a
combination of MO componentç due to nucleate boiling and Nm boiling proportional to the
fraction of wetted area, FA:
qlb=q1Fa + q v U - Fn) (6- 1)
where qi is the heat flux during the liquid contact and qv the heaî flux during the vapor contact.
An altemate approach assumes that the process king represented is ergodic, so that
instead of working in terms of an instantaneous, liquid contact-area fraction, the transition
boiling heat flux cm be written as
qlb=qrFe+ q v u -Fe) (6-2)
in ternis of a local iiquid contact-the fraction, Fe
Quantitative studies of liquid-solid contact in transition boiling regime have mostly used
either conductive probes or analysis of surface temperature fluctuations. Ragheb et al.
(1978,1979) used an insulated wire with a diarneter of 1 .O2 mm inserted through the wall to
measure the fraction of the wetted area. It was assumed that the wetted area fraction FA was zero
at the minimum film boiling point, FA =1 at the cntical heat flux point, and that the probe signal
varied linearly with FA in between.
By using an electncal conductance probe over a horizontal. flat, gold-plated copper
surface and studying contact phenornena in stable füm boiling of ethanol and water, Yao and
Henry (1 978) have confmed occurrence of Iiquid-solid contacts in pool boiling.
Lee et al. (1985) studied the wall temperature fluctuations by using a micro-
thennocouple flush-mounted on the boiling surface. Their results showed that the time-averaged
local liquid-contact fraction increased with decreasing surface superheat. The frequency of
liquid contact reached a maximum of -50 contactds at a surface superheat of -100 "K and
decreased gradually to 30 contactds near the critical heat flux.
Dhuga and Winterton (1985) developed a new method to detect liquid-solid contact by
measuring the impedance between a thin, electricaily insulating layer coated on the heating
surface and the boiling liquid in transition boiling. Using this technique, Rajabi and Winterton
(1988) found that the heat flwc during the liquid contact periods is not constant but f d s with
hcreasing surface temperature.
6.1.4. Liquid-solid contact frequency results for RI13
To obtain the liquid-solid contact frequency data for R113, quenching experirnents were
conducted on the ground using a data sampling rate of 900 Hz and the heat flux microsensor on
a copper disk as described in Chapter 2.
The magnitudes of heat flux and wall superheat fluctuations were obtained by
subtracting theü respective mean values from their instantaneous values, thus leaving ody the
flucniating components. The heat flux fluctuations and surface temperahue fluctuations are
show in Figures 6.1 and 6.2. The amplitude of heat flux fluctuations was of the order of several
hundred kw/rn2 while that of surface temperature fluctuations was less than SOC.
It is noted that the amplitudes of fluctuations in heat flux and surface temperature are
still quite large when the wall superheat has decreased even below the maximum heat flux point.
This is probably the fmt measurement to yield such heat flux and surface temperature
fluctuation data in quenching experiments. In steady boiling expenments, the local maximum in
heat flux is the Critical Heat Flux and the lower wail superheat region corresponds to nucleate
boiling regime in which the surface is presumably wetted and fluctuations in heat flux should be
quite small. However, Gaerfner (1965) photographically studied the high heat flux region in
nucleate boiling just below the Critical Heat Flux in pool boiling and observed that dry patches
appear in this region of the boiling curve. He called it as second transition region of nucleate
boiling regime. Whether the large heat flux fluctuations appearing in the present heat flux data
correspond to this second transition remains to be determined more clearly in the future.
Next, the frequency of direct liquid-soiid contact during transition boiling was obtained
from the power spectra of the surface temperature and heat flux data using the FFI' analysis
provided in Labtech Notebook software. For each spectrum, at least 512 data points were
processed over a given interval of time during which the surface temperature dropped by less
than about 10 OC. Typical power spectra for the surface temperature and heat flux data during
transition boiling are shown in Figure 6.3. For the temperature data, a dominant frequency was
found at about 10 Hz for many mns under different conditions. However, many dominant peak
frequencies were observed in the heat flux data, one at a low frequency of about 10 EIz and
others at higher frequencies that changed with wail superheat and fiow rate. The lack of high
frequency peaks in the temperature sensor data may be amibuted to the larger area of the
temperature sensor compared to the heat tlw; sensor and smaller sensitivity to temperature
R I 13
film boiling
1 transition boiling
I I I I 1 nucleate boiling
Time (second)
Figure 6.1. Typical magnitude of heat flux fluctuations for R113.
film boilingl . , 1 1 lnucleate boiling
O 2 4 6 8 10
Time (second)
Figure 6.2. Typicd magnitude of temperature fluctuations for R113.
f (Hz)
(a) Heat Flux
1.2e+5 ,
O 50 100 150 200
f (Hz)
(b) Surface Temperature
Figure 6.3. Typicai power spectra of q and Tw fluctuations for RI 13.
fluctuations. If the high frequency peaks are associated with smallxale rewet and dryout
events, the temperature sensor would not be able to capture these events, so it would show only
the low frequency peak.
In observing the boiling processes on the heat flux sensor captured by a high speed
video camera at a rate of 648 frames per second and 111OOOO second-' shutter speed it was
found that a large area encompassing the entire heat flux sensor could be cleariy seen when the
surface was rewetted More often, however, srnalier parts of the heat flux sensor could be seen
clearly which indicaies small-scale or partial wening of the surface. A clear view of the full
sensor was obtained about 8 to 10 times per second which corresponds to the low peak
frequency in the power spectrum data. Partially clear views of the sensor were obtained more
ofien that may correspond to the higher peak frequencies in transition boiling heat flux
fluctuations. It seems that rewetting and dryout of the surface involve different length and time
scales, with the large-scale rewet/dryout occurring less frequently than the small-scale rewet-
dryout phenornena.
Figure 6.4 shows the effects of wall superheat and flow rate on the Liquid-solid contact
fkequency detected by the present heat flux sensor. With decreasing wall superheat, the peak
frequency fmt increases and then decreases, so there exists a local maximum in the contact
frequency. Also, the curve at higher wall superheats shifts to higher wall superheats with
increasing flow rate. The dominant frequencies of about 50 to 70 Hz obtained in this work at
wall superheats of about 110 to 120 O C are somewhat higher than those of about 40 Hz at waU
superheats of about 40 O C for pool boiling of R- 113 with sirnilar liquid subcooling (Kalinin et
al., 1987).
6.13. Liquid-solid contact frequency results for PF5060
For the second coolant, PF5060, the quenching experiments were performed using the
heat flux sensor on a stainless steel disk and a data sampling rate of 500 Hz in both gravity
levels. Figures 6.5 and 6.6 show the heat flux data obtained at the same sampling rate and
similar flow rate under reduced and normal gravity during transition boiling over a short
penod of time in the same time period. The fluctuations in heat flux correspond to liquid and
vapor intemiittently covering the hot surface. Again, when the heat flux value was
significantly above zero, liquid wetted the hot surface and was being vaporized. Otherwise,
vapor covered the surface. The heat flux fluctuations in Figures 6.5 under reduced gravity
show that the liquid on the heat flux micro sensor gradually evaporated before vapor
occupied the sensor surface. However, in nomal gravity condition as s h o w in Figure 6.6,
the liquid evaporated about Nice as fast as that under reduced gravity before the sensor
surface was covered by vapor. It is apparent that the duration that the sensor surface was
covered by vapor is shorter under normal gravity than under reduced gravity.
Figure 6.4. The effects of wall superheat and mass flux on liquid-solid contact frequency for R 1 13.
RUN MG613
7.9 8.0 8.1 8.2 8.3 8.4 8.5
Time (second)
Figure 6.5. Heat flux fluctuations for PF5060 in transition boiling regime in p-g.
2500 RUN NG12
2000 -
CT
500 -
O -
14.7 14.8 14.9 15.0 15.1 15.2 15.3
Time (second)
Figure 6.6. Heat flux fluctuations for PF5060 in transition boiling regime in 1-g.
In order to analyze liquid-solid contact fiequency using an FFI' method provided in
SigmaPlot sofnvare, a treatment sirnilar to that applied to the raw data of R113 was also
perfomed for the data of PF5060. The heat flux fluchmtions are shown in Figure 6.7. Because
the magnitude of the surface temperature fluctuations was very small, the surface temperature
data could not be used in the Iiquid-solid contact frequency analysis.
The frequency of liquid-solid contact was again obtained from the power spectra of the
surface heat flux fluctuations and confiied by counting the number of heat flux peaks in a
given period of time. For each spectrum, 128 data points were processed over a given interval of
time during which the surface temperature dropped by 2 to 8 OC. A typical power spectrum of
the heat flux &ta during transition boiling is show in Figure 6.8.
Figures 6.9 to 6.11 show the effects of wall superheat and flow rate on the most probable
liquid-solid contact fiequency detected by the present heat flux sensor. Sirnilar to the RH3
results, the üquid-solid contact frequency curves for PF5060 also showed a local maximum in
the transition boiling regime, but the effect of flow rate on the contact frequency was small. The
curves at higher wall superheats shifted to higher wall superheats with increasing inlet
subcooling and gravity level (shown in Figures 6.12 and 6.13). In the nucleate boiling regime,
the iiquid-solid contact frequencies nearly remained constant ranging from 20 to 25 Hz under
both gravity conditions. The contact frequencies ranging from 20 to 40 Hz obtained under
normal gravity are higher than those ranging from 15 to 30 H i obtained under reduced gravity.
This is because gravity forces the Liquid to contact the surface whenever the vapor film collapses
due to insufficient heat transfer and vapor generation rates. The lack of gravity does not,
however, preclude liquid-solid contact possibly because of surîace tension effects, which can
spread the liquid film over the dry surface. Thus, there appear to be two modes of liquid-solid
contact, one due to vapor film collapse and the other due to the spreading of the liquid fdm.
Under normal gravity, both modes are equally significant and the liquid-solid contact fiequency
is considerably hi&. Under reduced gravity, the first mode becornes less important and the
contact frequency is reduced but remains sufficiently hi&, so that rewetting can proceed and the
boiling mode quickly changes to nucleate boiling.
Apparently, the liquid-solid contact frequency for R113 at higher wall superheats is
higher than that for PF5060. Sirnilar to the analysis of rewetting temperature and quench
velocity, lower vapor density and higher latent heat for RI 13 than those for PF5060 are believed
film boiling
-
transition boiling
8
RUN MG613
nucleate boiling
Time (second)
Figure 6.7. Typical magnitude of heat flux fluctuations for PF5060.
RUN MG613
1
i ~
Figure 6.8. Typical power spectra of heat fiux fluctuations for PF5060.
J
35 - 1414 kg/m2s p-g, A T , 35 OC 1247 kg/m2s
30 - 1 078 kg/m2s
h
N = 20 Y
Y-
15
nucleate boiling 1 transition boiling
Figure 6.9. Liquid-solid contact frequency for PF5060 in yg.
Figure 6.10. Liquid-solid contact frequency for PF5060 in 1-g with hiph inlet subcooling.
10 -
O
nucleate boiling ( transition boiling
I I l 1 1 I
nucleate boiling 1 transition boiling
Figure 6.1 1 . Liquid-solid contact frequency for PF5060 in 1-g for low inlet subcooling.
40 A p-g .1 O78 kglm2s, AT,,,= 35 OC
Figure 6.12. Cornparison of liquid-solid contact frequencies for PF5060 in p.-g and 1-g at low inlet flow rate.
45 A p-g , 1 4 1 4 kglm2s. AT,,,= 35 'C
Figure 6.13. Cornparison of liquid-solid contact frequencies for PF5060 in p-g and 1-g at high inlet flow rate.
to be responsible for the present results. Higher wall superheats for R113 in the transition
boiling regime could also contribute to higher liquid-solid contact frequencies.
6.2. Transition Boiling Heat Transfer of RI 13 As discussed in the last section, the liquid-solid contact frequency first increased
quickly to a maximum value and then gradually decreased with decreasing wall superheat in
the transition boiling regime. Also, Figures 4.1 and 4.6 showed that the instantaneous
maximum heat flux values generally increased with decreasing surface superheat in transition
boiling regime. If the heat flux for vapor covering penods during liquid-solid contacts is
neglected in equation (6.2), the heat flux in transition boiling calculated from equation (6.2)
would have a sharp increase and then a gradua1 increase with decreasing surface superheat as
shown in Figure 6.14. Therefore, two transition boiling regions, C and D, can be considered,
which correspond to the graduate and sharp increase in heat flux, respectively, in transition
boiling regirne (shown in Figure 6.14). A similar result has also k e n observed by Bamea and
Elias ( 1994).
Run NG-G5
Figure 6.14. Definitions of boiling heat transfer regions.
The effects of flow rate, inlet subcooling and gravity level are shown in Figures 6.15
and 6.16. With high subcooling and at normal gravity the flow rate affects transition boiling
more strongly than that in the absence of subcooling or gravity. It seems that the effect of
flow rate tends to diminish at higher flow rates in the absence of subcooling or gravity. These
coincide with Auracher's ( 1988, 1990) results for saturated and subcooled flow boiling of R-
114 and the results obtained by Cheng et al. (1978). In the near absence of subcooling and
gravity, the decreasing rewetting temperature induces a shift in the boiling curve to lower
wall superheats.
AIthough many researchers have experimentally and theoretically coriducted the
studies on transition boiling for several decades (Kalinin et al. (1987) and Auracher (1990)),
the mechanism of transition boiling is still unclear due to its very complicated processes.
This presents many difficulties for developing good models. However, many approaches
Figure 6.15. Transition boiling heat transfer for R 1 13 measured in I -g.
-
100
- -
1 O
- equation (6.6)
AT- = 25 O C O G=314kg/m%
G = 566 kg/rnZs A G = 1396 kg(m2s
saturated 0 G =377 kglrn2s a G = 705 kglrn2s O A a G = 1025 kg/m% v G = 1426 kg/m%
1 I
Figure 6.16. Cornparison of transition boiling heat transfer for R 1 13 in p-g and 1-g.
- - - - - - - equation (6.6)
p-g, ATs, = 25 O C
0 G = 549 kg/m2s 1 II G = 859 kg/m% - A G = 1308 kg/m*s
1 -g, ATçub = 25 O C
G=314kg/m2s A rn G = 566 kg/m% A G = 1396 kglmos
have been proposed for empirically correlating the experimental data The following form
is the simplest one and has been used by many researchers, such as Yilmaz and Westwater
( 1980),
where, a and b are constants.
Westwater ( 1989) denved a similar correlation for pool boiling of R- 1 13 on a copper
cy linder,
These two types of correlations work well for region C or D separately in Figure 6.14,
however, they do not fit the whole transition boiling curve. Thus, in order to correlate the
entire transition boiling data in the present work, Rohsenow's (1952) correlation,
was modified as follows,
where, a an b are constants.
The comparisons of the current data with equation (6.6) are shown in Figures 6.15
and 6.16, and good agreement has been achieved in most cases except near the quench front
region. The values of a and b are listed in Table 6.1. The value of b ranging from 2 to 5
increased with decreasing flow rate, while the value of a is very close to unity.
Table 6.1. The factors in correlation (6.6) for the mns in p-g and 1-g experiments.
NG-G2 237 21 3 14 1 .O70 4552 NGG5 248 22 566 1.098 3530
NG-GIO 288 25 1396 1.032 1.770
NG-GSS 200 44 377 1 .O29 3 .596 NG-GS6 200 45 705 1 -053 1.986 NG-GS7 202 46 1 025 0.988 1.820 NG-GS8 200 47 1426 0.980 1.81 1
Chapter 7
Maximum Heat Flux and Nucleate Boiling Heat Transfer
7.1. Maximum Heat FIux As discussed in the 1s t chapter, the vapor film collapse resulted in liquid-solid
contacts so that large fluctuations in heat flux and to a lesser extent in surface temperature
were seen at the onset of rewetting, which caused the surface temperature to decrease and
heat flux to increase sharply in a short time. The fluctuations in heat flux and surface
temperature revealed periods of rewetting and surface temperature recovery (dry-out)
occuning intermittently. At the beginning of the transition boiling period, the magnitude of
fluctuations in the surface temperature was the largest indicating that the dry-out period
lasted longer and rewetting was difficult to establish. With a further reduction in the surface
temperature, the liquid-solid contact frequency first increased to a maximum and then
decreased to a certain value. At the sarne time, heat flux first increased sharply and then
gradually reached a maximum, which is called the maximum heat flux.
The maximum heat flux usually marks the point of complete rewetting since partially
film boiling heat transfer can be neglected beyond this point. Therefore, the maximum heat
flux is normaily the boundary between the transition and nucleate boiling regimes which will
be discussed in the next section. In steady pool or flow boiling, the maximum heat flux is
usually called the Critical Heat Flux or CHF. In heat flux controlled boiling systems, if the
surface is heated beyond the temperature corresponding to the maximum heat flux, it could
result in a rapid surface temperature excursion and damage the boiling surface. Due to its
special importance in determining the operational Iimits of boilers, the CHF phenomena have
been studied for over six decades. however, the mechanisms responsible for the CHF
phenomena are still unclear at present. More CHF or maximum heat flux information is still
needed for the design of industrial devices, such as nuclear reactors, s t e m generators,
superconducting magnets and rocket engines.
Witte and Lienhard (1982) showed that the maximum heat flux in quenching
experiments might not be the sarne as those obtained during the steady boiling heat transfer
experiments. However, Ueda et al. (1983) showed that the critical heat flux coincided well
with the maximum heat flux obtained by their transient quenching tests with R113. Recently,
Huang et al. (1994) reported from their flow boiling experiments using water that the
transient effects at CHF point aimost disappeared except at very high m a s flux values, for
which the transient critical heat flux from quenching is somewhat Iower than the steady-state
value. In the present study, no steady-state flow boiling could be achieved because of
technical difficulties, so that a cornparison between steady and transient maximum heat flux
data could not be made. In the present study, q- will be used to represent the maximum heat
flux obtained from quenching experiments and qcw for the maximum heat flux From steady
state boiling expenments.
Figures 7.1 and 7.2 show the maximum heat flux data for R113 and PF5060 in
normal gravity and reduced gravity conditions, respectively. The maximum heat flux for
lower inlet flow rates increased with increasing flow rate, inlet subcooling, but decreased
with reduction in gravity. Some researchers, such as Katto (1980) and Westbye et ai. (1995)
have confirmed these trends. However, the effects of gravity and inlet subcooling for R113
diminished at higher mass fluxes. Sirnilarly, the maximum heat flux for PF5060 also
increased with increasing flow rate, inlet subcooling and gravity level. Unlike for R113, the
effect of liquid subcooling for PF5060 did not dirninish at higher m a s fluxes in the present
flow rate range. This difference could have been caused by higher vapor density and lower
latent heat of PF5060, because the liquid flow in the present range of flow rates may have
been unable to affect the vapor layer thickness much if a larger amount of vapor were
generated from the heating surface and the vapor film had large momentum. Mudawar and
Maddox (1989) and Willingharn and Mudawar (1992) studied flow boiling of dielectric
fluorocarbon (FC-72), which has similar themophysicai properties to that of PF5060, on one
or a linear array of discrete small heaters in a vertical rectangular channel. Their data showed
that CHF was less affected by increasing flow rate when the inlet velocity was less than 1.5
mis, which is consistent with the present data.
Some analyticd work in predicting C W has been performed for steady state pool
boiling and flow boiling. Katto (1985 and 1996) reviewed the advances in the study of CHI?
and tried to clarify the CHF mechanisrns developed to-date in different boiling systems, such
as pool boiling, steady state interna1 and external flow boiling. He pointed out that the
difficulties encountered in the study of CHF are in modeling the two-phase fluid behavior at
Figure 7.1. The maximum heat flux data for R113 in both gravity conditions.
Figure 7.2. The maximum heat flux data for PF5060 in p-g and 1 -g.
109
high heat fluxes near CHF, obtaining reliable measurements of the fluid state in the
immediate vicinity of the heater surface, and visudizing the key features associated with
boiling and CHF on the entire heater surface.
Two semi-theoretical models are very prominent in the literature, a hydrodynamic
instability model and a macrolayer dryout model. However, detailed examination of each
model is out of the scope of the present study. Thus, the present maximum heat flux data
were correlated by a dimensional analysis method, which is the most widely used in
correlating the experimental data.
The CHF phenornenon in saturated extemal flow or pool boiling is considered to be
associated with the hydrodynamic states, which can be represented in terms of the velocity,
force and length scdes, as follows, (i) the vapor velocity given by q,/pv hl" and liquid
velocity, u; (ii) the inertial forces (dependent on p, and pi), the viscous forces (dependent on
pv and pl), the surface tension CF, and the buoyancy force g(pr - p.); and (iii) the heater length,
L. Then, the dimensional analysis of those quantities yields a generd relationship as follows
(Kato, 1983):
For pool boiling, the terms relating to liquid velocity, U, are eliminated. For extemal flow
boiling, the term relating to buoyancy is eliminated. It is noticed that the second parameter is
the Weber number, We, and the last is the Reynolds number, Re. Because surface tension and
viscous forces are d l small compared to liquid momentum, it is difficult to decide which
dimensionless parameter cm be used or not. Cornmonly, the Weber number has been used in
correlating the experimentd data.
For intemal flow, Kato and Ohno (1984) developed the following relationship,
where d is tube diarneter and Ahi is the inlet subcooling enthalpy.
If saturated CHF data are correlated using the first two dimensionless parameters,
equation (7.2) can be written as follows,
where a, b and C are constants, and Weber number is given by We = G ~ L / cpl.
Katto and Kurata (1980) considered a submerged jet Bowing pardlel to a small
rectangular heater in normal gravity. The fluids tested were saturated water and RI 13 with
velocities ranging from 1.25 to 10 m/s. The CHF data were correlated in the form of equation
(7.3) and the values of a = 0.559, b = 0.264 and C = 0.186 were obtained. Using the Iiquid
and vapor densities of RI 13 in equation (7.3), the following equation can be obtained (note:
the parameter, 0 ' 1 / G ~ L , was used in their paper, which is the inverse Weber number, ~ e - ' ) ,
Yilmaz and Westwater (1980) obtained the following equation for flow boiling of
saturated RI13 over a horizontal copper tube at atmosphenc pressure in 1-g condition (note:
the vapor Weber number, Wev = ~~d /ap,, was used in their paper and, for convenience, their
equation has been converted to the fonn containing liquid Weber number),
From saturated RI 13 quenching data in the present study under normal gravity
conditions, the following equation was obtained.
It should be noted that the length of the plate, L, in the flow direction, not hydraulic
diameter, dh, was used in the Weber number in d l of the above correlations, because ordy the
bottom piate was heated in the present experiments.
The exponent of Weber number in equation (7.6) is close to those in equations (7.4)
and (7.5). However, the constant factor, 0.0339, is much higher than those in equations (7.4)
and (7.5). This is possibly caused by differences in flow boiling geometry (intemal flow vs.
external flow).
McGillis et al. (199 1) showed that their critical heat flux results determined from
subcooled R- 1 13 flow boiiing experiments (with subcooling ranging from 42 to 17 O C ) using
an array of simuiated rnicroelectronics devices on one wall of a vertical rectangular passage,
agreed well with a correlation developed by Mudawar and Maddox (1991), which is
functionally similar to equation (7.3). The exponent on Weber number in Mudawar and
Maddox's ( 199 1) correlation is -0.348.
The same equation forrn has been applied to the current maximum heat flux data for
subcooled inlet flow in both gravity conditions and the following correlations were obtained:
For R 1 13 in 1-g with subcooling of about 25 OC,
For R 1 13 in p-g with subcooling of about 25 OC.
For PF5060 in 1-g with subcooling of 18 OC,
For PF5060 in 1-g with subcooling of about 33 OC,
For PF5060 in p-g with subcooling of about 35 OC,
1 3 , saturated
q ~ ~ h k d . 0339 weOf.
Figure 7.3. The q,,,&Ghrv for RI 13 varies with We in both gravity conditions.
Figure 7.4. The q,,,&Ghrv for PF5060 varies with We in both gravity conditions.
113
The above correlations are plotted in Figures 7.3 and 7.4. It is noted that the
exponents on the Weber number in equations (7.7). (7.9) and (7.101, equal to -0.405, -0.355
and -0.395, respectively, are very close to -0.348 in Mudawar and Maddox's (1991)
correlation for subcooled inlet flow. Anotherinteresting result is that equation (7.6) is very
close to (7.8) for R113, which means that the effect of the absence of gravity on maximum
heat flux is very similar to the effect of the absence of subcooling in normal gravity. This
reveals that in the absence of gravity or subcooiing, liquid flow rate would strongly affect the
maximum heat flux.
More cornparisons of the maximum heat flux data for RI 13 and PF5060 wiil be
presented in the next chapter.
7.2. Nucleate Boiling Heat Trader of RI13
Of the three boiling heat transfer modes, nucleate boiling with its highly efficient heat
transfer and smaller wall superheat is the most important regime for industrial applications,
such as in boilers, cooling of rnicroelectronic chips, etc. Although a great amount of effort
ha been made so far to clarify the nucleate boiling rnechanism, the results have not been
fully satisfactory because of the complexity of the phenornena.
Conventionally or traditionally, the nucleate boiling regime on a quenching curve
starts at the maximum heat flux and extends down to the point of incipient boiling, as shown
in Figure 6.14. According to the dependence of heat flux on wall superheat obtained in this
work, there are also two regions in nucleate boiling, regions A and B. The time averaged heat
flux slowly decreased with decreasing wall temperature in region B and sharply decreased in
region A, which indicates that the bubble nucleation was likely suppressed readily due to the
smooth surface of the sensor.
It is noticed from Figures 4.1,4.6, 6.1 and 6.2 that the ampiinides of fluctuations in heat
flux and surface temperature were s a quite large when the wall superheat had decreased below
the maximum heat flux point. The analysis of heat flux fluctuations in the last chapter has
shown that the frequency of fluctuations in region B remained alrnost constant, at about 15 to 25
Hz, with decreasing wall superheat. This is probably the first measurement to yield such heat
flux and surface temperature data in quenching experiments.
In steady boiling experirnents, the local maximum in heat flux is the critical heat flux
and the lower wall superheat region corresponds to nucleate boiling regime in which the surface
is plesumably always wetted and fluctuations in heat flux are considered to be quite small. As
mentioned previously, Gaertner (1965) photographically studied the high heat flux region in
nucleate boiling below the critical heat flux in pool boiling and observed that dry patches appear
in this region of the boiling curve. He called it the second transition region of nucleate boiling
regime. From his study, the concept of macrolayer was established. The macrolayer mode1 was
applied to criticai heat flux analysis, for example, by Kano and Yokaya (1968), Yu and Mesler
(1977), Bhat el ai. (1983% 1983b 1986), Chyu (1987), Unai et al. (199 l), and Pasamehmetoglu
(1993). Whether the large heat flux fluctuations appearing in the present heat flux data
correspond to this second transition remain to be determined more clearly in future, since a
detailed analysis of the physical mechanisms in this region is out of the scope of the current
study.
Nucleate boiling heat transfer may be affected by many factors, such as wall
superheat, flow rate, liquid subcooling, gravity level, pressure, surface material and
roughness, and so on. The current experimental results for nucleate boiling in region B are
plotted in Figures 7.5 to 7.8. The heat flux in many cases increased with increasing flow rate
as shown in Figures 7.5 (a) to (c), which has been previously observed by many researchers,
such as Yilmaz and Westwater (1980). Also, as mentioned in Chapter 4, boiling curves shifi
to lower wall superheats under low subcooling or reduced gravity conditions as shown in
Figures 7.6 and 7.7. Merte and Clark (1964) and Zell et al. (1989) in pool boiling
experiments and Westbye et al. (1995) in quenching experiments al1 found that the effect of
gravity level on nucleate boiling is srnall. The current results coincide with their results in the
overlapped region as evident in Figure 7.7. However, it is found that the nucleate boiling
region B under low subcooling or low gravity conditions is wider than those in the high
subcooling and normal gravity case, with about 30 O C difference in the wall superheat. Figure
7.6 (b) shows the strong effect of subcooling on nucleate boiling. Again the behavior in the
absence of subcooling or gravity is very similar, as shown in Figure 7.8. This may be caused
by an increase in the thermal boundary layer thickness, which is caused by the accumulation
of vapor bubbles under these conditions. A thicker thermal boundq layer was observed by
Ervin et al. (1992) in pool boiling experiments under microgravity. This is not fully clear and
needs to be confirmed in further studies.
Figure 7.5 (a). The effect of 80w on nucleate boiling in 1 -g with subcooled inlet.
Figure 7.5 (b). The effect of flow on nucleate boiling in 1-g with saturated inlet.
1000 1-9, saturated
equatlon (7.1 2)
100
---O-- G = 377 kg/m% --O-- G = 502 kg/m2s ---a--- G = 705 kglmzs
. - --V--- G = 1025 kg/m% Q' --O-- G = 1426 kglmzs
1 I 1 1
Cr-9, AT*ub = 25 O C - (cl equation (7.12)
e
---O- G = 549 kg/m% ---o.- G = 773 kglm2s
a - ---A-+- G = 1303 k-s -
Figure 7.5 (c). The effect of flow on nucleate boiling in p-g with subcooled inlet.
Figure 7.6 (a). The effect of subcooling on nucleate boiling in 1-g.
Figure 7.6 (b). The effect of subcooling on nucleate boiling in p-g.
20 40 60 80 100 120 140
T, - T,, ec)
Figure 7.7. The effect of gravity on nucleate boiling heat transfer.
Figure 7.8. Cornparison of nucleate boiling in the absence of subcooling and gravity.
Many theoretical and experimental snidies on nucleate boiling either in pool boiling
or flow boiling heat transfer have been conducted and many models have been reported. A
simple form of correlation cornrnonly used to predict nucleate boiling heat transfer is given
by
4&= &, -Ta (7.12)
where a and b are empirical constants chosen to fit both pool and flow boiling data.
The correlations for the present data are shown in Figures 7.5 to 7.8 by solid lines and
good agreement has been obtained for most of the data The values of constants, a and b, are
listed in Table 7.1.
It is found that for most of the runs, as the flow rate increases, the constant, a,
decreases while the exponent, b, increases. This coincides with the increase in CHF as
discussed in Section 7.1. Table 7.2 lists the values of b obtained by previous researchers, and
it shows that the exponeat, b, has a value of about 3 for pool boiling, which is much higher
than the present values shown in Table 7.1. On the other hand the value of b obtained by
Yilmaz and Westwater (1980) for flow boiling in normal gravity ranged from 1.5 to 2 in
rough agreement with the present data, however, their flow velocities were much higher than
those in the present experiments. A further study is required to explain the differences among
the previous and present sets of 1 -g data.
Table 7.1. The constants in equation (7.12) for p-g and 1-g experiments.
MG-D23 207 22 549 0.55 1.43 MG-D22 205 22 773 O. 15 1.74 MG-D20 213 21 1258 0.70 1.48 MG-D 15 192 20 1303 O. 17 1.81 MG-D2 1 1 94 22 1330 0.0020 2.77
MG-D8 180 13 617 0.23 1.72 MG-D 12 177 16 746 0.13 1.86
NG-G2 237 21 333 9 .O3 0.8 1 NG-GO 220 20 474 0.023 2.12 NG-G 1 222 22 578 0.6 1 1.45 NG-G 10 288 25 1396 0.98 1.39
Table 7.2. The data of b obtained by other researchers.
Researchers Type of boiling b
Rohsenow ( 1952) pool boiling Borishansky (1969) pool boiling
S tephan ( 1980) pool boiling for refngerants Clarke ( 1963) pool boiling for cryogens
Gaertner ( 1965) pool boiling 0.6 3.72 (O m/s)
Yilmaz and Westwater flow boiling of low subcooled R- L 13 1.96 (2.4 m/s) (1980) outside a horizontal copper tube 1.57 (4.0 mis)
1 -49 (6.8 m/s)
Chapter 8
Cornparison of RI13 and PFS060 Resuits
The quench velocity, rewetting temperature and maximum heat flux are very
important parameters in a quenching process, which involves high surface temperatures and
may be therrnodynamically or hydrodynamicdly controlled. ln this section, the quenching
data obtained with R113. PF5060 and other fluids with different thermophysicai properties,
will be compared to evaluate effects of the fluid properties on quenching heat transfer both in
normal and reduced gravity. The thermophysical properties of RI 13 and PF5060 are listed in
Table 8.1.
Table 8.1. Fluid and thermophysical properties of R- 1 13 and PF5060.
Boiling Curve, Quench Velocity and Rewening Temperature
Figure 8.1 compares the boiling curves for RI13 and PF5060 obtained durhg
quenching at sirnilar mass flow rates and inlet subcooling in 1-g and p-g conditions. The
boiling curves shift to higher wall superheats with increasing inlet subcooling and gravity
level for both fluids. It is apparent that the boiling curve for RI 13 shifts to higher wall
superheats because the rewetting temperature is higher for RI13 than PFSO6O.
Cornparisons of rewetting temperature and quench velocity between the two fluids
have been discussed in Chapter 5. In summary, the rewetting superheat and quench velocity
for R113 were much higher than those of PF5060. This is mainly due to a lower vapor-to-
liquid density ratio of Ri 13 (about 40% smaller than that of PF5060), and higher latent heat
(about 40% greater than that of PF5060). Because a liquid having a larger latent heat will be
evaporated less per unit heat input, a smaller amount of vapor will be generated near the hot
surface. Also, the smaller the vapor density is, the smaller the vapor momentum would be to
push the liquid away from the hot surface. Both of these factors caused the vapor film of
R113 to collapse at higher wail superheats and the hot surface to quench faster compared to
PF5060. Although surface tension is also expected to affect the quenching and boiling heat
transfer characteristics, the effect is believed to be smaller than those of latent heat and vapor
density, as discussed in detail in Appendix I(3).
Figure 8.1. Cornparison of the boiling curves of R 1 13 and PFS060.
Maximum Heat Flux for Subcooled Inlet Flow
The maximum heat flux data for subcooled inlet flow of RI 13 and PF5060 in 1-g and
p.-g are shown in Figures 8.2 and 8.3. In 1-g condition, the maximum heat fluxes for R113
with inlet subcooling of 25 OC fa11 between those for PF5060 with inlet subcooling of 33 OC
and 18 O C . In reduced gravity, the maximum heat fluxes for RI 13 with inlet subcooling of 25
O C are only slightly greater than those for PF5060 with inlet subcooling of 35 OC. The
maximum heat flux data shown in Figures 8.2 and 8.3 can be best fit by the following
equations, which has the same fonn as equation (7.2) and are shown in Figures8.4 and 8.5.
A PF5060,l -g,~T-=33~C
P FSO6O.l -g,ATw=l 8OC - equation (8.1)
100 1 I
Figure 8.2. Comparison of maximum heat fluxes for subcooled inlet flow in 1-g.
joo t m R113.p-g,~~~,=25~C
A P ~ 5 0 6 0 , p - g ~ ~ ~ ~ ~ = 3 5 ~ ~ - equation (8.2)
Figure 8.3. Cornparison of maximum heat fluxes for subcooled inlet flow in p-g.
For 1-g and inlet subcooling ranging frorn 18 OC to 33 O C for both fluids,
For p-g and inlet subcooling of 25 OC for R113 and 35 OC for PF5060.
where the liquid Jakob number is defined as Ja, = AT,& CP, hl"
The exponent on the density ratio, pv/pi, in 1-g is 1.034, which is higher than the
values for saîurated extemal flow boiling discussed in Chapter 7. This exponent value could,
however, underestimate the effect of the density ratio, pJpi, on the maximum heat flux, so it
should be modified in future work.
Figure 8.4. Comparison of subcooled q&Ghiv vs. We with the prediction in 1-g.
Figure 8.5. Comparison of subcooled q 2 G h l v vs. We with the prediction in p.-g.
Conclusions and Recommendations
An experimental apparatus has been designed and built incorporating micro heat flux
and surface temperature sensors, and used to determine the effects of gravity as well as liquid
flow rate and subcooling on quenching of a hot horizontal surface. A condenser especidy
designed for reduced gravity experiments aboard the NASA's KC-135 and DC-9 parabolic
aircraft could supply single-phase liquid to the test section under rnicrogravity conditions.
The quenching heat transfer results have been analyzed and compared with the existing data
obtained by other researchen. The following conclusions c m be drawn from the results of the
present work.
1. Instantaneous fluctuations of large amplitude in heat flux and surface temperature
have been detected from the onset of rewetting to high wall-superheat nucleate boiling in
both gravity conditions. The data clearly showed the liquid-solid contacts occurred not only
in transition boiling regime but also in nucleate boiling at hi& wall superheats in both gravity
conditions.
2. The boiling curves shifted to higher wall superheats with increasing liquid
subcooling and gravity level. The boiling curves for RI13 were obtained in a higher wall
superheat range compared to those for PF5060, because of higher rewetting temperature for
RI 13 than that for PF5060.
3. The quench velocity and rewetting temperature decreased for R113 but only
showed very slight decreases for PF5060 in reduced gravity. The quench velocity and
rewetting temperature for RI13 were higher than those for PF5060 mainly due to higher
latent heat, higher surface tension and smaller vapor density for R 1 13.
4. A peak frequency was found in the liquid-solid contact frequency curves for R113
in normal gravity and PF5060 in both gravity conditions. The contact frequency for PF5060 at
higher wall superheats increased with increasing liquid subcooling and gravity level.
5. The maximum heat flux increased with increasing flow rate, but decreased with
reduction in inlet liquid subcooling and gravity level except for R 1 13 at high flow rates.
6. The rewetting temperature, transition and nucleate boiling heat transfer and the
maximum heat flux for RI 13 showed good similarity between the mns with saturated liquid
injection in normal gravity and subcooled liquid injection with subcooling of 25 OC in
reduced gravi ty.
7. The nucleate boiling regime of RI13 covered a wider range of wdl superheat
below the maximum heat flux in the absence of gravity or subcooling than in the case of high
inlet subcooling and in normal gravity. The results showed a suong effect of subcooling on
nucleate boiling in reduced gravity.
Future experirnental work on quenching of a hot surface in both gravity conditions
should be conducted to obtain more data for film boiling heat transfer and quench velocity.
The liquid-solid contact phenornena need to be investigated more by using real-time data
acquisition combined with real-time high speed video photography, so that the boiling
mechanism at high wall superheats can be understood. The condenser should be rnodified to
overcome the effects of low frequency g-jitter in future parabolic flight experirnents.
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Appendix 1 (1)
The Effect of the HFMS Disk Material
In this study, a copper HFMS disk was used in the rewetting experiments using RI 13
and a stainless steel disk for PF5060. In order to determine the effect of HFMS disk materid
on quenching and boiling heat transfer characteristics, two quenching mns were performed
under normal gravity using RI13 and the stainless steel HFMS disk. The main difference in
the properties of copper and stainless steel is in thermal conductivity. At 100 OC - 200 OC, the
thermal conductivities are 377 W/m°C for copper and 18 W/m°C for stainless steel (Welty,
1984). The rewetting ternperanire, quench velocity and maximum heat flux obtained using
the copper and stainless steel HFMS disks are compared below.
For the stainless steel HFMS disk, the experimental conditions are listed in Table A. 1.
Table A. 1. The experimental conditions for R 1 13 on stainless steel disk.
S.S. disk T w (OC) Tb (Oc) Uin (m/s) G (kg/mzs) Run #1 23 1 22 0.642 969 Run #2 239 23 0.884 1335
For the copper EIFMS disk, the corresponding results had to be estimated from the
data presented in the main text.
Rewetting temperature
For the copper HFMS disk, the rewetting temperatures were obtained from equation
(5.8) which best fit the data with the inlet liquid subcooling of 25 O C and the initial wall
ternperature ranging from 2 15 OC to 288 O C .
Table A.2. Cornparison of rewetting ternperature between two HFMS disks.
G Copper disk S tainless steel disk (S.S. I Cu) 100% ocg/m2s, (OC) (OC>
969 182 169 93% 1335 187 176 94%
The difference in the rewetting temperatures is seen to be quite small, and the effect
of the HFMS material is not significant.
Quench velocity
It is hard to directly compare the quench velocity data because of a lack of data for
similar flow rate and initial wall temperature conditions. It is found from Figure A. 1 that the
quench velocities for the inlet velocity ranging from 0.125 mls to 0.917 m/s, the initial wall
temperature ranging from 235 OC to 253 OC and the inlet liquid subcooling of 25 OC showed a
linear variation. By performing a linear regression analysis, the following equation was found
to best fit those data.
In the following table, the quench velocity for the copper disk was calculated by
equation (A. 1 ).
Table A.3. Comparison of quench velocity on two HFMS disks
h Copper disk Stainless steel disk (S.S. / Cu) 100% W s (mm/s (mm+'s O. 642 13.3 7 -3 55% 0.884 17.1 L 0.0 58%
The quench velocities for the stainless steel disk are estimated to be about half of
those for the copper disk. This is reasonable because of lower thermal conductivity of
stainless steel, which makes heat removal from the disk more time consuming and thus,
slows down the quench velocity.
Maximum heat flux
For the copper disk, the maximum heat fluxes were calculated from equation (7.7)
which could best fit the data for inlet Iiquid subcooling of 25 OC.
Table A.4. Comparison of the maximum heat flux on two H F ' S disks.
G Copper disk Stainless steel disk (S.S. / Cu) 100% (kplm2s) (kwlm') (kw/m2) 969 676 654 97% 1335 7 19 804 112%
The differences are small and the effect of the sensor material (copper and stainless
steel) on the maximum heat flux is also srnail.
- R I 13, Copper disk - AT,, =25 O C
T,,, = 235 - 253 O C
data - U, = 15.76U, + 3.1 7
1 r
Figure A. 1 . The quench velocity for a copper disk HFMS.
III
The Effect of g-jitter on Quenching Characteristics and Boiling Heat Transfer
Using FFT, the typicai power spectra of gravity level fluctuations in KC-135 and DC-
9 aircraft were obtained as s h o w in Figures A.2 and A.3. A broad band spectrum ranging
from O to 50 Hz was found aboard the KC-135 but the dominant frequency of about 0.3 Hz
was found aboard DC-9 (Note: the peak at 60 Hz is believed to be caused by electrical field).
Comparing the liquid-solid contact frequency data of PF5060. ranging from 17 to 28 Hz,
with the dominant frequency of g-jitter (0.3 Hz) on DC-9, there does not appear to be any
effect of g-jitter. Aiso, qualitatively comparing the data on boiling curve, rewetting
temperature, quench velocity and the maximum heat flux in both gravity conditions, the
features are sirnilar. So, the same conclusion can be drawn.
Figure A.2. The typicai power spectrum of gravity level fluctuations in KC-135.
Figure A.3. The typical power spectrum of gravity level fluctuations in DC-9.
v
Appendix I (3)
The Effect of Surface Tension
The effect of surface tension on quenching and boiling heat transfer characteristics
can be qualitatively determined by examining the following aspects, Critical Velocity for
Kelvin-Helmholtz instability, bubble growth rate, Cntical Heat Flux and minimum film
boiling temperature in pool boiling.
Critical veIocity
From the analysis of Kelvin-Helmholtz instability, the critical velocity can be derived
as follows (Careys, 1992),
The critical velocity is proportional to the one fourth power of surface tension, Uc - 0'". This means that the stability of a liquid-vapor interface with higher surface tension can
be maintained at higher liquid and vapor relative velocities. If the Critical Heat Flux is
proportional to the critical velocity as in the hydrodynamic theory of CHF (Zuber, L959), the
Critical Heat Flux could be expressed as follows.
In this correlation, C is constant and
Therefore, the strongest effect of thermophysical properties on the Critical Heat Flux
cornes from the latent heat, hi,, and to a lesser extend, vapor density, pv. The effect of surface
tension, a, is believed to be srnaller than the first two.
Critical Heat Flux in pool boiling
Zuber (1959) developed an analytical model for Critical Heat Flux in pooi boiling.
The following equation was developed by Lienhard and Dhir based on Zuber's model.
In the correlation, 1/4 1/4
q c - hiv pvl" pi Q
The same conclusion as in the discussion of critical velocity can be drawn.
Bubble growth rate
At high wail superheats in the present study, the bubble growth is inertia controlled
and the Rayleigh equation describes the bubble growth rate.
The bubble growth rate decreases with increasing surface tension and liquid density.
If the surface tension is small and the bubble growth is at the beginning stage (& cc R), the
bubble growth has little dependence on surface tension.
Minimum film boiling temperature in pool boiling
Berenson (1961) obtained a correlation based on Rayleigh-Taylor instability in pool
film boiling, given by equation (5.5) as follows.
In the correlation,
A L - h v pv pl 9 6 #2
The minimum film boiling temperature is mainly affected by latent heat, vapor
density and liquid density. The effect of surface tension is srnalier. Higher surface tension
induces higher rewetting temperature because more vapor is needed to maintain the stability
of the Iiquid-vapor interface at higher waI1 temperatures.
Summary
In surnrnary, surface tension plays a role in stabilizing the liquid-vapor interface or
restonng the perturbed interface. That means. the higher the surface tension is, more vapor
Bow is needed to destabilize the interface. Also, for higher surface tension, the bubble
growth rate and inertia would be smaller. For RL 13 with about twice greater surface tension
than that for PF5060, the rewetting temperature and quench velocity would be higher than
those for PF5060.
However, the effect of surface tension can be considered to be smaller than that of
latent heat and vapor density. The main effects of thermodynamic properties of Buid on the
quenching heat transfer characteristics corne from the latent heat and the vapor density.
Appendix II
Uncertainty Analysis
In order to assess the reliability of the experimental results and the effects of inlet
flow rate, liquid subcooling and gravity level on quenching characteristics and boiling heat
transfer, the uncertainties on heat flux, surface temperature and quench velocity have k e n
anaiyzed and are presented as follows.
Heat Flux
The accuracy of HFMS was 35% of the reading and that of two amplifiers kO.546. So,
the accuracy of heat flux measurement was caiculated from equation (A. 10) to be *.O% of
the reading.
(A. 10)
The maximum tirne-averaged heat flux in the present study was less than 900 kwlm2.
Then, the maximum heat flux uncertainty was estimated to be k45 kw/m2.
2. Surface Temperature
The surface temperature was calibrated by the manufacture and re-checked before
conducting experiments. The maximum surface temperature uncertainty was estimated to be
A52 OC.
3. Quench Velocity
As descnbed in Chapter 5, the quench velocity was caiculated by dividing the
distance between two heat flux sensors (7 mm) by the time difference, (tz - ti), between the
minimum heat fiuxes at quench point measured by two heat flux sensors.
(A. 1 1)
The uncertainty in the distance between two heat flux sensors was esthated to be
MSmm or the accuracy was &7% of 7 mm. The uncertainty in the time at which liquid starts
to rewei the hot surface was estimated to be M . O 1 second for LOO Hz sampling rate and
M.002 second for 500 Hz. Then, the uncertainty in the time difference between two sensors
was M.02 second and M.004 second for 100 Hz and 500 Hz data sampling rate,
respectively. So. the accuracy was estirnated to be 10% for the minimum time difference of
0.2 second for R113 with 100 Hz sampling rate and 0.2% for the minimum time difference of
2.0 second for PF5060 with 500 Hz sampling rate. Therefore, the accuracy of quench
velocity was estimated to be 13% for R 1 13 data and 7% for PF5060 data as calculated from
equation (A. 1 2).
