CONVECTIVE HEAT TRANSFER OF BINARY MIXTURES
UNDER FLOW BOILING CONDITIONS
E. V. McAssey Jr.,
Villanova University, Villanova, PA USA
S. G. Kandlikar
Rochester Institute of Technology, Rochester, NY USA
Abstract
The present study represents part of an effort in considering the replacement of ethylene-glycol/water mixtures with propyleneglycol/water mixtures for engine cooling in automotive application. Experimental results are presented for the heat transfer
coeffic ient under flow boiling conditions for water and mixtures of water with ethylene-glycol and propylene-glycol. The
range covered includes the single-phase non-boiling region through the fully developed subcooled flow boiling region to
saturated boiling. Since there are no predictive methods available in literature for subcooled flow boiling of binary mixtures, a
preliminary comparison with pure component correlations is presented in the fully developed boiling region. Further work is
being undertaken to extend the methodology to the partial boiling region and the significant void region.
NOMENCLATURE
α
= heat transfer coefficient, W/m2 K
Bo
= boiling number =
q ''
G i fg
= convection number , ( ρ G / ρ L )0 .5 ((1 − x) / x )0 .8
= specific heat of liquid, J/kgK
= mutual diffusion coefficient of 1 in 2 in liquid
phase, m2 /s
dT/dx = slope of the bubble point curve for the mixture
FFl
= fluid surface parameter
G
= mass flux, kg/m2 s
∆h LG = latent heat of vaporization, J/kg
q ′′ = heat flux, W/m2
r
= radius, m
T
= temperature, K
T1,2 = measured temperatures at radial positions, r1 and r2
Co
cp,L
D12
1/2
V1
x1
y1
c  k  dT

= volatility parameter, V1 = P.L 
( x1 − y1 )
∆hLG  D12  dx1
= mass fraction of more volatile component in
liquid phase
= mass fraction of more volatile component in
vapor phase
Subscripts
B
= binary
CBD = convective boiling dominant
conv = convective component
f
fl
LO
nb
NBD
sat
TP
w, wall
=
=
=
=
=
=
=
=
bulk fluid
fluid
liquid only
nucleate boiling component
nucleate boiling dominant
saturation value
two phase
wall
INTRODUCTION
The objective of this paper is to present experimental
data on flow boiling from the subcooled boiling region to the
saturated boiling region. Using this data base, various
predictive schemes will be evaluated. The data includes single
and binary component mixtures.
With the greater emphasis on performance, there has
been an increased interest in achieving higher heat transfer
coefficients by the use of boiling heat transfer. An area of
considerable activity is low-pressure flow boiling. Internal
combustion engines are cooled with ethylene-glycol/water
mixtures operating at approximately two atmospheres and
encompassing the heat transfer regimes from single phase
through saturated boiling. Advanced nuclear reactor systems
provide emergency core cooling in the pressure range of one
to two atmospheres. Computer manufacturers are involved in
investigations using boiling heat transfer to cool high
dissipation components.
Although a large body of work exists for saturated
boiling, subcooled boiling under flow conditions requires
additional investigation. The purpose of the present study is to
present experimental results obtained for binary mixtures
under operating conditions representative of engine cooling
systems. The broader purpose of the experiments is to
compare the thermal performance of aqueous mixtures of
propylene-glycol and ethylene-glycol. Although ethyleneglycol/water mixtures have performed satisfactorily in internal
combustion engines, leaks from automotive cooling systems
are responsible for accidental human and animal exposures.
Ingestion of ethylene-glycol can be harmful or fatal even in
relatively small doses. Propylene-glycol is less toxic than
ethylene-glycol and possesses very similar heat transfer
characteristics, and, therefore, appears to be an ideal
replacement. The resulting boiling heat transfer data base can
provide a resource to evaluate heat transfer coefficient
prediction methods.
In addition to the mixture data, a limited amount of
data for water was also obtained. The test conditions provide
results from single -phase flow, sub-cooled boiling, and
saturated boiling.
The resulting data base contains
approximately 7500 test points. Although most of this data
was obtained at a fixed mixture concentration of 50/50 by
volume, a limited amount of data was obtained for 70/30 and
30/70 aqueous mixtures for both coolants.
BACKGROUND
Experimental data for flow boiling with mixtures under
subcooled conditions is limited. A number of studies are
reported in literature for flow boiling of mixtures under
saturated conditions. Among the predictive methods available
in literature for saturated flow boiling of mixtures, those
presented by Jung [1], Bennett and Chen [2], and Kandlikar
[3] are noteworthy. Jung [1] presented extensive results for
refrigerant mixtures and based upon this data developed a
correlation scheme involving 25 constants. Bennett and Chen
[2] developed a correlation based upon the widely used Chen
[4] correlation for saturated boiling. Kandlikar [3] extended
his theoretical model developed for pool boiling of binary
mixtures (Kandilkar [5] ) to flow boiling and compared his
correlation with available data for refrigerant mixtures with
good agreement.
Although ethylene-glycol/water mixtures have been
used as engine coolants for over fifty years, there is very little
heat transfer data available in the open literature. In an
automobile cooling system, the working fluid is generally a
mixture of water and either ethylene-glycol or propyleneglycol. The normal mixture concentration is 50/50 by volume.
Finlay et al. [6] presented experimental results for an ethyleneglycol/water mixture covering an operating range appropriate
to automotive engine cooling conditions. Most of the data
were obtained under constant pressure conditions using a
copper test section. Some data were also obtained for other
test section materials including cast iron and aluminum, and
for constant flow-rate operation. These results showed
reasonable agreement with analytical predictions based upon
the Chen correlation at low surface heat fluxes. However, at
higher fluxes under subcooled boiling conditions, the same
model tended to under-predict the surface temperature.
McAssey, Stinson, and Gollin [7] presented test results
comparing propylene-glycol/water and ethylene-glycol/water
mixtures for a range of conditions similar to those existing in
normal engine operation. For the range of test conditions, the
resulting data spanned the spectrum from single-phase
convection to saturated boiling.
These investigators
concluded that the overall performances for both coolant
mixtures were very similar. This paper also presented
comp arisons between analytical predictions, again based on
the Chen correlation, and experimental results. In general, the
analytical results under-predicted the surface temperature
when the difference between the surface temperature and the
fluid temperature exceeded 60 o C.
Bhowmick, Branchi, McAssey, and Gollin [8]
presented additional data on both fluid mixtures for a wider
range of conditions. In their work, the inlet velocity was
varied from approximately 0.4 m/s to 2.5 m/s and the surface
heat fluxes reached a maximum of 1.8 MW/m2 . In addition,
comparisons were presented between experimental results and
analytical predictions based upon the Chen correlation. The
experimental results showed that both mixtures had similar
thermal performances. In general, the analytical results underpredicted the measured wall temperature by a significant
margin.
Bhowmick et al. [9] presented comparisons between
predicted results and experimental data for water and aqueous
mixtures of ethylene-glycol and propylene-glycol. Although
the Chen [4] correlation provided the best approximation to
the experimental results, the method tended to over-predict the
heat transfer coefficient. Using water data obtained as part of
the engine test program, a revised Chen correlation was
developed. The revision involved a modification to the S
factor in the Chen correlation. With this revised correlation,
the authors showed improved predictions for both ethyleneglycol and propylene-glycol mixtures.
Kandlikar [10] presented a method to predict subcooled
boiling of a pure component. It was based on the saturated
flow boiling correlation developed by Kandlikar [11]. Three
regions were identified under subcooled flow boiling after the
onset of nucleate boiling (ONB): partial boiling region, fullydeveloped boiling region, and significant void-flow region.
The criteria for identifying each region along with the
correlations for specific regions were also presented. These
correlations showed good agreement with published
experimental data for water and several refrigerants.
In the present work, Kandlikar's [10] model for
subcooled flow boiling in the fully developed boiling region is
extended to the binary mixtures. The onset of nucleate boiling
and partial boiling region will be covered in the follow up
work of this investigation.
DESCRIPTION OF TEST FACILITY
Flow Loop
The flow loop consisted of a test section, pump, accumulator
tank, rejection heat exchanger, and required piping. Figure 1
presents a schematic layout of the test loop. For this program,
the loop was operated in the controlled flow mode. In this
mode of operation, the test section is provided a constant
volumetric flow-rate under all operating conditions. Since
most of the pressure drop occurs across the control valve,
changes in the test section pressure drop due to heating have a
small effect on the flow-rate to the test section. The bypass
line around the test section was installed to allow operation on
the constant-head portion of the pump performance curve and
to reduce the vapor content of the fluid entering the heat
exchanger. Table 1 indentifies the types of instruments used
in the experiment. The primary flow-rate instrument was a
Measurement
Temperature
Pressure
Pressure Drop
Test Section Flowrate
Loop Flow-rate
Power Input
Test Section Fluid
Temperature
Table 1 Instrument Accuracy
Type of Instrument
E-type thermocouple
Pressure gauge
Diffe rential pressure transducer
Turbine flowmeter
Rotometer
Voltage and current measurement
E-type thermocouples
turbine flow meter (FL-1) located upstream of the test section.
Test section inlet pressure and temperature were measured
by a pressure gauge (P-1) and a thermocouple (RTD-1)
respectively. A similar arrangement (P-2 and RTD-2) located
downstream of the test section was used to measure exit
conditions.
Accuracy
+/- 1.2 o C
+/- 1% F.S.
+/- 0.25% F.S.
+/- 0.50% of reading
+/- 0.75% F.S.
+/- 1% of reading
+/- 1.2 o C
produced a constant head source for the loop pump, which was
a centrifugal pump capable of producing 0.3 m3 /min. at 400
kPa. A control valve located upstream of the test section was
used to provide the required flow to the test section. After
leaving the test section, the bypass and the test section flows
combined before entering the heat exchanger.
Test Section
The test section (shown in Fig. 2) was constructed from
a 114 mm diameter aluminum cylinder with a length of 165
mm. The horizontal flow channel consisted of a 9.53 mm
diameter hole drilled along the centerline of the cylinder. The
test section was heated by ten cartridge heaters located on a
97.9 mm diameter circle. Each heater was capable of
producing 1,000 watts at 120 volts. By operating at slightly
higher voltages, it was possible to dissipate approximately 11
Figure 2 Test section
Figure 1 Schematic of the flow loop
The accumulator tank was used to maintain the system
pressure. Pressure control was accomplished by means of
adjusting the liquid level or the argon pressure. An argon
blanket was maintained in the top of the accumulator tank and
the pressure in this region was adjusted using an external
argon supply. The tank also contained a heater to control the
test section inlet temperature. The accumulator tank also
kW in the test section.
The test section was connected to the flow loop by
means of inlet and exit calming sections, which have the same
inside diameter as the test section. The upstream calming
length and the test section each had an L/D ratio of 17:1. This
ratio was considered to be sufficient to provide fully
developed flow to the heated length. In order to minimize heat
loss from the system, the entire flow loop and test section
were insulated. The insulation on the test section consisted of
over 200 mm of fiberglass insulation.
The test section temperatures were measured by 20 Etype thermocouples, mounted at five axial stations along the
cylinder (see Fig. 2). Using two thermocouples on the same
radial line, the surface temperature can be calculated using Eq.
(1).
ln(r0 / r1)
ln( r2 / r1 )
2.5e+6
(1)
All thermocouple measurements were made in the
same horizontal plane. This arrangement provided two
essentially equal surface temperature measurements at each
axial location.
Surface Heat Flux - W/m 2
Twall = T1 + ( T2 − T1 )
bias error is, therefore, ± 12%.
In developing these data, a reference test was
established and repeated several times. Figure 6 presents data
for propylene-glycol/water obtained over a two month period
Instrumentation
The instrumentation for this program was chosen to
determine various test parameters and test section
temperatures. Test section flow-rates were measured with
turbine flowmeters and rotometers. The rotometers were
calibrated using a weigh tank with the mixtures at the required
operating temperature of 85 o C. The turbine flowmeters were
then calibrated against the rotometers. A high temperature
thermocouple calibrator was used to calibrate each sensor over
the complete operating range. Table 1 presents an estimate of
the overall accuracy of the instruments used in these tests.
2.0e+6
Ethylene-glycol/water
x/L = 0.345
1.5e+6
x/L = 0.5
x/L = 0.654
1.0e+6
5.0e+5
0.0
80
100
120
140
160
Wall Temperature - oC
180
200
Figure 4 Surface heat flux versus wall temperature for
ethylene-glycol/water mixture at an inlet
velocity = 1.33 m/s, Tinlet = 85 o C, Pexit = 205 kPa
Surface Heat Flux - W/m2
2.5e+6
∆(Twall-Tfluid)/(Twall-Tfluid)
50
Propylene-glycol/water
x/L = 0.345
x/L = 0.5
x/L = 0.654
2.0e+6
1.5e+6
1.0e+6
5.0e+5
40
positive deviation
30
negative deviation
20
10
0
-10
-20
-30
-40
-50
0.0
0.0
100
120
140
160
180
200
5.0e+5 1.0e+6 1.5e+6 2.0e+6
Surface Heat Flux - W/m2
2.5e+6
220
Figure 5 Comparison of measurement error in (Twall - Tfluid)
for both mixtures at an inlet velocity = 1.3 m/s,
o
exit pressure = 205 kPa, inlet temperature = 85 C
Since the present study involves the prediction of
heat transfer coefficient, the uncertainty in the measurement of
this quantity must be examined. The experimental test
program was designed to simulate normal engine operating
conditions, and, therefore, the flu xes and wall temperatures are
quite high. Figures 3 and 4 present surface heat flux versus
measured wall temperature for a typical set of test conditions
for the binary mixtures. The two parameters determining the
heat transfer coeffcient are the measured power and the
temperature difference (Twall –Tfluid ). The power measurement
bias error is ± 2%. Using the ± 1.2 o C uncertainty in
temperature measure-ment, Fig. 5 shows that the uncertainty
in (Twall – Tfluid) varies with flux. In the range of interest
above 250,000 W/m2 flux, the bias error is ± 10%. The total
Heat Transfer Coefficient - W/m 2K
Wall Temperature - oC
Figure 3 Surface heat flux versus wall temperature for
propylene-glycol/water mixture at an inlet
velocity = 1.33 m/s, Tinlet = 85 oC and Pexit = 205 kPa
The data represent results from
three different tests at the same nominal
conditions over a two month period.
Dashed line represent 95% confidence
level
30000
25000
20000
15000
10000
5000
0
0.0
5.0e+5
1.0e+6 1.5e+6 2.0e+6 2.5e+6
Surface Heat Flux - W/m2
Figure 6 Comparison of heat transfer coefficient for
propylene-glycol/water for an inlet velocity = 1.3 m/s,
o
exit pressure = 205 kPa, inlet temperature = 85 C
at the same nominal test conditions. This figure shows that
most of the data falls within or very close to the 95%
confidence level. Using these results as the precision error,
the overall uncertainty in heat transfer coefficient is ± 13%.
FFl in eqs. (3) and (4) is a fluid-surface parameter
related to the nucleation characteristics which is taken as the
mass fraction averaged value for the pure components. The
single-phase heat transfer coefficient
α LO is obtained from
the Petukhov and Popov [12] and Gnielinski [13] correlations.
For further details, refer to Kandlikar [3].
DATA BASE
The experimental results were obtained for a range of
conditions representative of internal combustion engine
cooling system conditions for both normal and abnormal
operation.
The normal aqueous equilibrium mixture
recommended by engine manufacturers for both ethyleneglycol and propylene-glycol is 50/50 by volume. However, it
was
recognized
that
this
mixture
concentration
recommendation was not necessarily followed by the end user.
Therefore, tests were performed at two other concentrations to
examine the effect on coolant performance. In addition to the
two binary mixtures, experiments were conducted with 100%
water.
DISCUSSION
Region II
Moderate diffusion-induced suppression region,
0.03< V1 ≤0.2, and Bo >10-4 , includes the region where mass
diffusion effects begin to affect the heat transfer. Here, the
nucleate boiling is suppressed and heat transfer is well
represented by the following correlation developed for the
CBD (Convective Boiling Dominant) region given by eq. (5).
α TP. B = α T P.CBD
(5)
Region III
Severe diffusion-induced suppression region, (a) for
0.03<V1 ≤0.02 and Bo<10-4 , and (b) V1 >0.2;
The theoretical approach presented here combines the
saturated flow boiling model for mixtures by Kandlikar [3]
and subcooled flow boiling for pure components by Kandlikar
[10]. As stated earlier, only the fully developed boiling region
is investigated.
Saturated flow boiling heat transfer with binary
mixtures was classified by Kandlikar [3] into three regions
depending on the level of suppression as defined by a
volatility parameter, V1 .
The mass diffusion effects further reduce the heat transfer
given by the CBD equation, eq. (4), through a mass diffusion
factor FD applied to the nucleate boiling component as seen in
eq. (7).
Region I
Near azeotrope region, V1 ≤0.03, is applicable where
the effects of mass diffusion are insignificant, as is the case
with azeotropic mixtures.
  c

FD = 0.6781+  P .L
  ∆hLG

α TP. B
α
R
= larger of S
Tα
TP , B , NBD
(2)
TP , B , CBD
α TP, B ,NBD and α TP, B ,CBD are given by eqs. (3) and (4)
respectively using mixtures properties.
α TP .B = 1136
. Co −0 .9 (1 − x ) 0.8 α LO
+ 667.2 Bo 0.7 (1 - x ) 0.8 FFlα LO
1/ 2
 κ 


 D 
 12 
dT 
x1,s − y1,s
dx1 

(
0.7
α NT ,B = 0.6683Co
(1 − x ) α LO
0.8
+ 1058.0 Bo 0.7 (1 − x )0 .8 FFl α LO
α TP ,CBD = 1136
. Co
−0 .9
(3)
0.7
(1 − x ) α LO
+ 667.2 Bo 0.7 (1 − x ) 0 .8 FFl α LO
(4)
(7)
(Tw − Tsat )
( Tw − Tf )
(8)
In Eq. (8), the coefficient 1058 corresponds to the nucleate
boiling term in the nucleate boiling dominant equation given
by eq. (3). Kandlikar and Bulut [14] extended the model
α TP = 667.2 Bo FFl α LO
0.8
−1
Instead of using the interface concentrations x1,s and y1,s , the
equilibrium liquid and vapor concentrations may be used in
eq. (7) for simplicity. In the fully developed region of
subcooled boiling, Kandlikar [10] presented the following
correlation, which is rewritten in Eq. (8) in terms of the wall
minus fluid temperature difference.
α TP = 1058.0 Bo FFl α LO
−0. 2
)
(6)
(Tw − Tsat )
(Tw − Tf )
(9)
presented in eqs. (2)-(7) to the fully developed region of
subcooled boiling with mixtures as given by the following
equation.
Equation (9) is expected to apply in the moderate
diffusion-induced suppression region. An additional factor FD
would be applicable where diffusion-induced suppression is
severe.
50000
Prediction G = 1235 kg/m2s
Prediction G = 2375 kg/m2s
Prediction G = 411 kg/m2s
2
25000
20000
Single Phase
Prediction
15000
G = 2540 kkg/m 2s
10000
5000
G = 1321 kg/m2s
G = 440 kg/m2s
0
-40 -30 -20 -10 0 10 20 30 40 50 60 70 80 90 100
Measured Wall Superheat - o C
40000
Figure 9 Comparison between equation (9) correlation and
experimental data for ethylene-glycol/water
mixture at Tinlet = 85 o C, and P exit = 205 kPa, x/L = 0.5
35000
30000
25000
Single Phase
Prediction
20000
G = 2375 kg/m2 s
G = 1235 kg/m 2 s
10000
0
-40
-20
0
20
40
60
Figure 7 Comparison between equation (9) correlation and
experimental data for water at Tinlet = 85 oC and
Pexit = 205 kPa, x/L = 0.5
35000
Heat Transfer Coefficient - W/m2K
G = 411 kg/m2 s
5000
Measured Wall Superheat - oC
Predicted Heat Transfer Coefficient- W/m2K
30000
45000
15000
Prediction G = 2540 kg/m2s
Prediction G =1321 kg/m2s
Prediction G = 440 kg/m2s
35000
Heat Transfer Coefficient - W/m K
Experimental,
Experimental,
Experimental,
Experiment,
Experiment,
Experiment,
40000
35000
Prediction G = 2540 kg/m s
Experiment,
Prediction G = 1321 kg/m s
Experiment,
Prediction G = 440 kg/m s
2
2
25000
20000
15000
10000
5000
Single Phase
Prediction
G = 2540 kg/m2 s
G = 1321 kg/m2 s
x/L = 0.5
G = 440 kg/m2 s
0
-40 -30 -20 -10 0 10 20 30 40 50 60 70 80 90 100
Measured Wall Superheat - oC
30000
25000
2
Experiment,
30000
+20%
Figure 10 Comparison between equation (9) correlation and
experimental data for ethylene-glycol/water
mixture at Tinlet = 85 oC, and Pexit = 129 kPa
-20%
20000
15000
10000
x/L = 0.5
5000
0
0
5000 10000 15000 20000 25000 30000 35000
Experimental Heat Transfer Coefficient - W/m 2K
Figure 8 Comparison between equation (9) and experimental heat
transfer coefficient for water at T inlet = 85 oC, Pexit = 205 kPa
Figures 9 through 12 show comparisons between
prediction, in the fully developed boiling region, from Eq. (9)
and experimental data for two binary mixtures, ethyleneglycol/water and propylene-glycol/water. Each mixture had
the same equilibrium concentration of 50/50 by volume. In
general, the predictions based upon Eq. (9) provided fair
agreement for the ethylene aqueous mixture. In Figs. 11 and
12 for the propylene-glycol/water mixture, it can be seen that
the approach significantly under-predicts the heat transfer
coefficient when compared to the experimental data. Several
factors must be considered. The fluid-surface parameter, Ffl,
Heat Transfer Coefficient - W/m2K
Heat Transfer Coefficient - W/m2K
The present work falls under moderate diffusion
induced suppression region as indicated by Kandlikar and
Murat [14]. Figure 7 presents a comparison for pure water
data with predictions based upon Eq. (9) with the convection
dominant coefficient. In the fully developed region, the
agreement is reasonable. Figure 8 presents a comparison
between prediction and experiment for all flow rates. Except
for a few points, all the data could be predicted within ± 20%.
Note that some of the data points fall under partial boiling
region where a more detailed scheme presented by Kandlikar
[10] is applicable.
35000
30000
25000
Experimental,
Experimental,
Experimental,
Prediction G = 2391 kg/m2 s
Prediction G = 1243 kg/m2s
Prediction G = 414 kg/m2 s
x/L = 0.5
20000
15000
Single Phase
Prediction
10000
G = 2391 kg/m2 s
G = 1243 kg/m2 s
5000
x/L = 0.5
G = 414 kg/m2 s
0
-40 -30 -20 -10 0 10 20 30 40 50 60 70 80 90 100
Measured Wall Superheat - oC
Figure 11 Comparison between equation (9) correlation and
experimental data for propylene-glycol/water
mixture for Tinlet = 85 oC, and P exit = 205 kPa
Heat Transfer Coefficient - W/m 2K
was taken as unity, which may not be true for this mixture.
Also, the mass diffusion effects may not be as severe, and a
constant between 667 and 1058 may be appropriate. Work is
continuing on the calculation of diffusion suppression factors.
Figure 13 presents a comparison of prediction and experiment
for all the ethylene-glycol/water data at 129 kPa and 205 kPa.
Except for the partial boiling region, the agreement is very
good, within ± 25%.
Experiment,
Experiment,
Experiment,
35000
30000
Prediction G = 2391 kg/m2 s
Prediction G = 414 kg/m2s
Prediction G = 1243 kg/m2s
25000
20000
15000
10000
Single Phase
Prediction
G = 2391 kg/m2s
x/L = 0.5
5000 G = 1243 kg/m2 s
G = 414 kg/m2 s
0
-40 -30 -20 -10 0 10 20 30 40 50 60 70 80 90 100
Measured Wall Superheat -oC
Predicted Heat Transfer Coefficient - W/m2K
Figure 12 Comparison between equation (9) correlation
and experimental data for propylene-glycol/water
mixture at Tinlet = 85 oC, and Pexit = 129 kPa
35000
Pexit = 129 kPa
Pexit = 205 kPa
30000
fully developed boiling correlation, Eq. (9).
Further
improvement is expected by incorporating specific
correlations given by Kandlikar [10] in the partial boiling
region.
Equation (9) with an Ffl of unity was also shown to
yield good results for a 50/50 by volume mixture of ethyleneglycol/water. For the propylene-glycol/water mixture, this
approach tended to under-predict the heat transfer coefficient
over a range of operating conditions. Estimation of accurate
fluid-surface parameter, and further evaluation of diffusion
effects for this mixture are expected to improve the results.
Finally, additional operating conditions and mixture
concentrations are being investigated.
ACKNOWLEDGEMENTS
The authors wish to express their appreciation to the
ARCO Chemical Co. and the Cummins Engine Co who
provided financial support for the experimental phase of this
investigation.
REFERENCES
1. Jung, D.S., 1988, “Horizontal Flow Boiling Heat Transfer
Using Refrigerant Mixtures,” Ph.D. Dissertation,
University of Maryland.
2.
Bennett, D.L., and Chen, J.C., 1980, “Forced Convective
Boiling in Vertical Tubes for Saturated Pure Components
and Binary Mixtures,” AICHE J., Vol. 26, No. 3, pp. 454461.
3. Kandlikar, S.G., 1998, “Boiling Heat Transfer with Binary
Mixtures: Part I- Pool Boiling,” ASME Journal of Heat
Transfer, Vol. 120, No. 2, pp. 380-387.
+ 25%
25000
-25%
20000
3.
15000
10000
5000
0
0
5000 10000 15000 20000 25000 30000 35000
Chen, J.C., 1966, “A Correlation for Boiling Heat Transfer
to Saturated Fluids in Convective Flow,” Industrial and
Engineering Chemistry, Process Design and Development,
Vol. 5, No. 3 pp. 322-329.
5. Kandlikar, S.G., 1998, “Boiling Heat Transfer with Binary
Mixtures: Part II- Flow Boiling in Plain Tubes,” ASME
Journal of Heat Transfer, Vol. 120, No. 2, pp. 388-394.
Experimental Heat Transfer Coefficient - W/m 2K
Figure 13 Comparison between equation (9) and experimental
heat transfer coefficient for ethylene-glycol/water
mixture at T inlet = 85 oC
CONCLUSIONS
Experimental results are obtained for subcooled flow
boiling of water and aqueous mixtures of ethylene-glycol and
propylene-glycol under a range of conditions generally
encountered in automotive engine-cooling applications. The
experimental results are compared with the existing predictive
methods. Since there are no methods available for subcooled
flow boiling of mixtures, a preliminary comparison is carried
out following the model developed by Kandlikar and Mulut
[14], which combines the methodology presented by
Kandlikar [10] for subcooled boiling of pure components and
by Kandlikar [3] for saturated flow boiling of mixtures.
The pure water data could be predicted within 30
percent accuracy using the single-phase correlation and the
6.
Finlay, I.C., Boyle, R.J., Pirault, J.P., and Biddulph, T.,
1987, 'Nucleate and film Boiling of Engine Coolants
Flowing in a Uniformly Heated Duct of Small Cross
Section," SAE Technical Paper Series 870032.
7. McAssey, E. V., Stinson, C., and Gollin, M., 1995,
“Evaluation of Engine Coolants Under Flow Boiling
Conditions,” Proceedings of the ASME Heat Transfer
Division, HTD-Vol. 317-1, p. 193-200.
8.
Bhowmick, S., Branchi, C., McAssey, E.V., and Gollin,
M., 1996, “Heat Transfer Performance of Engine
Coolants Under Subcooled Boiling Conditions,” ASME
ICE-Vol. 26-2, 1996 Spring Technical Conference.
9. Bhowmick, S., Branchi, C. McAssey, E.V., Gollin, M. and
Cozzone, G., 1997, “Prediction of Heat Transfer in Engine
Cooling Systems,” Proceedings of the 4th World
Conference on Experimental Heat Transfer, Fluid
Mechanics, and Thermodynamics, Brussels, Belgium.
10. Kandlikar, S.G., 1998c, “Heat Transfer Characteristics in
Partial Boiling, Fully Developed Boiling, and Significant
Void Flow Regions of Subcooled Flow Boiling,” ASME
Journal of Heat Transfer, Vol. 120, No. 2, pp. 395-401.
11. Kandlikar, S.G., 1990, “A Generalized Correlation for
Saturated Two-Phase Flow Boiling Heat Transfer Inside
Horizontal and Vertical Tubes,” ASME Journal of Heat
Transfer, Vol. 112, pp. 219-228.
12. Petukhov,B. S., and Popov, V. N., 1963, “Theoretical
Calculation of Heat Exchange in Turbulent Flow in Tubes
of an Incompressible Fluid With Variable Physical
Properties,” High Temperature Heat Physics, Vol. 1, No.
1, pp. 69-83.
13. Gnielinski, V., 1976, "New Equations for heat and Mass
Transfer in Turbulent Pipe and Channel Flow,"
International Chemical Engineer, Vol. 16, pp. 359-368.
14. Kandlikar, S.G., and Bulut, M., 1999, “An Experimental
Investigation On Flow Boiling Of Ethylene-Glycol/ Water
Mixture,” Paper submitted for presentation at the 1999
ASME National Heat Transfer Conference, Albuquerque