International Journal of Refrigeration
34 (2011) 1846-1853

Boiling heat transfer of HFO-1234yf flowing in a smooth small-diameter
horizontal tube
Shizuo Saitoha＊, Chaobin Dangb, Yoshitaka Nakamurab, Eiji HiharabaDepartment of Mechanical Engineering, The University of Tokyo, 7-3-1, Bunkyo-ku, Tokyo
113-8656, Japanb
Institute of Environmental Studies, Graduate School of Frontier Sciences, The University ofTokyo, 5-1-5 Kashiwanoha, Kashiwa-shi, Chiba, 277-8563, Japan*Corresponding author. E-mail: [saitoh@hee.k.u-tokyo.ac.jp] Tel./Fax: [+81 3 5841 6325]
Abstract
The flow boiling heat transfer coefficient of the low-GWP (global warming potential)refrigerant HFO-1234yf inside a smooth small-diameter horizontal tube (inner diameter: 2mm) was experimentally investigated. The local heat transfer coefficient was measured at heat
fluxes of 6–24 kW m-2, mass fluxes of 100–400 kg m-2 s-1, evaporating temperature of 288.15
K, and inlet vapor quality of 0–0.25. The results show that the effect of heat flux on the heattransfer was large at low vapor quality, while the effect of mass flux was large at high vapor
quality. The heat transfer coefficient of HFO-1234yf was almost the same as that of R-134a.
The heat transfer coefficients calculated based on correlations with Saitoh et al. agreed wellwith the measured values compared to other correlations. The measured pressure drop agreed
well with that predicted by the Lockhart-Martinelli correlation.
Keyword: Boiling heat transfer, Pressure drop, HFO-1234yf, Small-Diameter tube
1. Introduction
The release of the MAC (mobile air conditioning) directive by the EU, which bans the use ofrefrigerants with global warming potential (GWP) above 150 in new types of mobile air
conditioning from 2011 in the EU market, has triggered the research and development in asearch for new refrigerants. Carbon dioxide is considered to be a promising candidate.
Recently, HFO-1234yf, which is another promising candidate, was jointly developed by
Honeywell and DuPont. Because the GWP of HFO-1234yf is as low as 4 and its
thermophysical properties are similar to those of R-134a, it is expected to be a drop-in
solution for current mobile air conditioners. Several experimental studies have been
conducted on its thermophysical properties and cycle performance to estimate the feasibilityof using this new refrigerant in mobile air conditioners. Increasing concern due toenvironmental predictions has led to the reconsideration of refrigerants in other applications.
-1-One approach is to use a refrigerant mixture of HFO-1234yf + HFC-32 to obtain a highsystem coefficient of performance (COP). The thermophysical properties of the refrigerant
mixture are also being evaluated (i.e., Arakawa et al. 2010).
The system performance of an actual heat pump is lower than that of the theoretical cycle
because of heat transfer loss inside the heat exchangers and pressure drop along the duct.
Therefore, the heat transfer performance must be analyzed to evaluate the system performanceof an actual heat pump system when a new refrigerant is considered as well as when
designing heat exchangers. In this study, the boiling heat transfer of the refrigerant
HFO-1234yf flowing in a smooth small-diameter horizontal tube (inner diameter (ID): 2 mm)
was experimentally investigated. The measured local heat transfer coefficient of HFO-1234yf
was compared with that of R-134a, and a prediction method for the evaporation heat transfer
coefficient of HFO-1234yf is discussed.
2. Experimental apparatus and procedure
Figure 1 shows a schematic of the experimental system used to measure the heat transfer
coefficient and pressure drop and observe the flow patterns of HFO-1234yf. The test loop
includes a Coriolis-type flow meter, refrigerant temperature controller, flow control valve, test
tube (evaporator), and sight glass. The purity of the HFO-1234yf used was over 99.7%. To
reduce the heat loss from the test tube to the environment, the entire test tube was placed
inside an air duct with the air temperature controlled to be equal to the evaporating
temperature. The flow rate and inlet pressure of HFO-1234yf were controlled by adjusting thefrequency of the magnetic gear pump and opening of the flow control valve. The vapor
quality at the inlet of the test evaporator was adjusted by the amount of heat supplied to therefrigerant in the pre-heater. A sub tank was used to adjust the amount of refrigerant in the test
loop. When the performance of the condenser was poor, the evaporation pressure wascontrolled by the amount of refrigerant. Figure 2 shows the measurement points for outersurface temperature and pressure with the 2 mm ID test tube, and thermocouples attached tothe tube. Table 1 lists the specification of the test tube and measurement interval of
temperature and pressure. The test tube was heated by direct electrification using a DC power
supply connected to two electrodes soldered at the flanges of the two ends of the test tube.
The pressures of the refrigerant in the test tube were measured using a precision aneroid
manometer. The temperatures of the outer surface of the test tube were measured midway
between the top and bottom of the tube along the surface using T-type thermocouples (outer
diameter (OD): 0.1 mm), and the temperatures of the inner wall of the tube were calculated
from the measured temperatures of the outer wall of the tube using Fourier’s law. An 8 mthick Teflon sheet was inserted between each thermocouple and the test tube to prevent thecurrent from affecting the thermocouples.
All the thermocouples were calibrated by using a high-precision platinum resistancethermometer sensor (Chino, Model CNA) with an accuracy of ±0.03 K. The accuracy of the-2-

calibrated thermocouples was within ±0.1 K. The mass flow rate (and thus the mass flux G)
was measured by using the Coriolis-type flow meter (Oval, Model E010S-IN-200) with anaccuracy of ±0.1%. The pressure was measured using the precision aneroid manometer(Nagano Keiki, Model NKS) with an accuracy of ±1.5 kPa. The electrical input power wasmeasured using a voltmeter and ammeter to confirm that the heat generated by directelectrification was transferred well to the fluid and that the heat gain from surroundings waswithin 3%. The experimental conditions are summarized in Table 2. To compare accumulated
data of boiling heat transfer for R-134a, the evaporation temperature is chosen to be 288.15K
as used in previous research for R-134a. The local heat transfer coefficient hexp in the test tube
was determined using the following equation:q
hexp 
(1)
Twall  Tsat
where Twall is the temperature of the inner wall and Tsat is the saturation temperature at the
local refrigerant pressure calculated by interpolation of the adjoined pressure gauges. All
experimental data were collected after the steady state was reached for temperature, pressure,and refrigerant flow.
In this study, the saturated vapor pressure of HFO-1234yf was correlated using experimental
data from Tanaka and Higashi (2010), Nicola et al. (2010) and Hulse et al. (2009);
thermodynamic properties at the saturation state were calculated using Akasaka et al.’s
method (Akasaka et al., 2010). Thermal conductivities of liquid and vapor were calculatedusing Latini et al. and Chung et al.’s methods, respectively (Poling et al., 2001). Liquid andvapor viscosities were determined using Hulse’s et al. correlation (2009) and theChapman-Enskog equation (Poling et al., 2001), respectively. The properties of the refrigerant
R-134a were calculated using REFPROP version 8.0 (Lemmon et al., 2007).
3. Results and discussion
3.1 Flow boiling heat transfer coefficient
3.1.1 Effect of heat flux
Figure 3 shows the variation in the heat transfer coefficient against the vapor quality. The
mass flux was kept at 200 kg m-2 s-1; we compared the measured results for three different
heat fluxes: 6, 12, and 24 kW m-2. At the lowest heat flux of 6 kW m-2, the measured heattransfer coefficient increased with the vapor quality, showing that the convective heat transfer
intensifies with increasing quality. The dryout quality was about 0.8 and did not change withheat flux. Increasing the heat flux from 6 kW m-2 to 12 and 24 kW m-2 showed that the heattransfer coefficient increases with heat flux at low vapor quality; thus, nucleate boiling is thedominant heat transfer coefficient mechanism at low vapor quality.
3.1.2 Effect of mass flux
Figure 4 shows the effect of mass flux on the boiling heat transfer at heat flux of 12 kW m-2.
The dryout occurs at vapor quality of 0.8 for all the conditions. In the high-quality region-3-

(>0.4), the heat transfer coefficients at both mass fluxes (200 and 400 kg m-2 s-1) increased
with increasing vapor quality, and the heat transfer coefficient was higher at 400 kg m-2 s-1
than at 200 kg m-2 s-1. At a mass flux of 100 kg m-2 s-1, the effect of vapor quality on the heattransfer coefficient was weak. The results suggest that in the high vapor quality region, forced
convective evaporation is dominant.
3.1.3 Comparison between HFO-1234yf and R-134a
Figure 5 shows a comparison between the boiling heat transfer performances of HFO-1234yf
and R-134a at a mass flux of 300 kg m-2 s-1 and heat flux of 12 kW m-2. The figure shows thatin the wide vapor quality region, the difference between the heat transfer coefficients of
HFO-1234yf and R-134a is small, which may be because the differences in their
thermodynamic properties are small. Saturation properties of HFO-1234yf and R-134a attemperature of 288.15K are shown in table 3. The properties of HFO-1234yf and R-134a arecalculated following Brown et al.(2009) and using REFPROP ver.8, respectively. The gas
density, latent heat, liquid thermal conductivity, and surface tension of HFO-1234yf and
R-134a at 288.15 K were 26.3 and 23.76 kg m-3, 156.5 and 186.59 kJ kg-1, 0.0724 and 0.0854
W m-1 K-1, and 0.0077 and 0.0094 N m-1, respectively. In the following, the heat transfer
coefficient of HFO-1234yf and that of R-134a is compared using the prediction model byconsidering the contribution of nucleate boiling and convective heat transfer. The heattransfer coefficient due to nucleate boiling is evaluated by Stephan-Abdelsalam(1980) correlation, and that of forced convection for liquid alone in the tube 2 mm
ID is evaluated by Dittus-Boelter equation. At a temperature of 288.15K, massflux of 300 kg m-2 s-1 and heat flux of 12 kW m-2, the heat transfer coefficients(HTCs) of nucleate boiling and forced convection for HFO-1234yf are 2.44 kW m-2
K-1 and 0.89 kW m-2 K-1, respectively, and that for R-134a are 2.30 kW m-2 K-1 and
0.92 kW m-2 K-1. The difference between the HTCs of HFO-1234yf and that of
R-134a is small.
3.1.4 Boiling heat transfer coefficient vs. Lochart-Martinelli parameter
In general, the flow boiling heat transfer in a tube is considered to be a combination of
nucleate boiling heat transfer and forced convective evaporation. In forced convective
evaporation, flow boiling data can be correlated with the form hexp  hL 1 X  , where hL isnthe heat transfer coefficient of the liquid alone and X is the Lockhart-Martinelli parameter.
When the superficial liquid Reynolds number, Rel, is smaller than 1000, flow in the liquidphase is laminar, and hL  4.36 lD. When the Rel is larger than 1000, liquid flow is

turbulent and the hL can be calculated by Dittus-Boelter equation. The magnitude of forced
convective evaporation can be expressed by the gradient n of a linear regression of theexperimental data. Figure 6 shows the measured heat transfer coefficients in the pre-dryout
region plotted against 1/X. Most of the present data can be fitted to the regression line-4-

hexp / hL  1 X 

0.77. The scattering of the data is larger at low 1/X (<7) than at high 1/X. At

low 1/X, i.e., low vapor quality, both nucleate boiling and forced convective evaporation
occur, whereas at high 1/X (>7), the scattering of the data is small because forced convective
evaporation is dominant. When flow in the liquid phase is laminar, the heat transfer
coefficient is barely influenced by the Lockhart-Martinelli parameter.
3.1.5 Pre-dryout heat transfer and predictions of some correlations
Saitoh et al. (2005, 2007) experimentally studied the boiling flow heat transfer mechanism of
R-134a in tubes with ID of 0.51, 1.12, and 3.1 mm; the Chen-type correlation was modified
by considering the effect of the tube diameter characterized by the Weber number in the gas
phase.
For the pre-dryout heat transfer, the experimental data and predictions of some correlations
were compared. These included Saitoh et al.’s, Yoshida et al.’s (1994), Kandlikar’s (1990),
and Gungor-Winterton (Gungar and Winterton, 1987) correlations. Details on these
correlations are shown in Table 4. The properties of the refrigerant HFO-1234yf were
calculated using the results reported by Brown et al. (2009). Figures 7(a–d) show acomparison between the experimental heat transfer coefficient hexp and calculated heattransfer coefficient hcal based on our modified Chen-type correlation (Saitoh et al., 2007) withothers for HFO-1234yf in the 2 mm ID smooth horizontal tubes. hcal based on the Saitoh et al.
correlation agreed well with hexp. However, hcal calculated based on the Yoshida, Kandlikar,
and Gungor-Winterton correlations did not agree well with hexp. Figure 7(e) and (f) show thecomparison of four correlations and measured heat transfer coefficient against vapor quality.
In the pre-dryout region, the predicted values by all the four correlations increase
with vapor quality. In the low quality region (x<0.5), the predicted values of
Saitoh et al., Kandlikar and Gugor-Winterton are close to the measured values
except for that of Yoshida. In the high quality region (0.5<x<0.75), the differenceof the predicted values of Kandlikar correlation and Gungor-Winterton correlation
with measured results are large. The predicted value by the correlation of Saitoh
et al., which was proposed for flow boiling heat transfer of R-134a, approximately
coincides with the measured value of HTC for HFO-1234yf. Table 5 lists the mean
deviation and accuracy (defined as the fraction of data within ±20% error) for each of thethree correlations. The Saitoh et al. correlation showed an improved mean deviation andaccuracy (9.2% and 92.8%, respectively) compared to the Yoshida (20.4% and 51.9%),
Kandlikar (18.7% and 51.9%), and Gungor-Winterton (15.5% and 66.9%) correlations.
3.2 Pressure drop
The measured two-phase pressure drop ( p ) was compared with that predicted results using
the Lockhart - Martinelli correlation, which defines the pressure drop as.

-5- dp  dp 
 dp        l2     g2
 dz 
 dz  l dz  g(2)

where l and g are the two-phase multipliers in the liquid and gas phases, respectively. They
are defined as l2  1  c / X  1/ X 2 and g2  1  cX  X 2 (Chisholm, 1967), where X is the
Lockhart-Martinelli parameter, which is the square root of the ratio between the pressure drop
assuming liquid flow alone and assuming gas flow alone. When flows of liquid and gas
phases are turbulent, c = 20; when liquid phase is laminar and gas phase is turbulent, c = 12.
In this study, the flow condition of the refrigerant was determined at the inlet of the evaporator.
Figures 8 shows the measured pressure drops and those predicted using the
Lockhart-Martinelli correlation. The measured pressure drops agreed well with that predictedusing the Lockhart-Martinelli correlation.
4. Conclusions
The flow boiling heat transfer of the refrigerant HFO-1234yf in a small-diameter horizontal
tube was experimentally investigated. The local heat transfer coefficients were measured at
mass fluxes of 100–400 kg m-2 s-1, heat fluxes of 6–24 kW m-2, and an evaporating
temperature 288.15 K. The results are summarized as follows.
1. At low vapor quality, nucleate boiling heat transfer is the dominant heat transfer
mechanism, and at high vapor quality, forced convective evaporation is dominant.
2. The boiling heat transfer coefficient of HFO-1234yf is almost the same as that of R-134a.
3. The heat transfer coefficient predicted by the Saitoh et al. correlation in the pre-dryout
region agrees with the measured data within a range of ±20%, while the measured pressure
drops agreed well with that predicted using the Lockhart-Martinelli correlation.
NomenclatureBoboiling numberCgfriction factor for gas, Cg = 0.046
Cl
friction factor for liquid, Cl = 16
cpl
specific heat at constant pressure in liquid phase, J kg-1 K-1
D
inner diameter of a tube, mGmass flux, kg m-2s-1
gacceleration of gravity, m s-2
hexp
experimental boiling heat transfer coefficient, W m-2 K-1
hL
liquid phase heat transfer coefficient, W m-2 K-1
0.8

1/ 3

 GD c  
Rel  1000, hL  0.023 l  l   pl l D  l   l 
Rel  1000, hL 

4.36l
D
-6-

p
q
Re
Re l 

Twall
Tsat
We
x
X

pressure, Pa
heat flux, W m-2
Reynolds number
Gl D

l

, Re g Gg Dginside-wall temperature, K
saturation temperature, K
Weber number
vapor quality
Lockhart-Martinelli parameter
Re l  1000, Re g  1000
1 x 
X 

 x 

0.9

 g   l 

0.5

 l 
 
 
 g

0.1

Rel  1000, Reg  1000
0.5

0.5

C 
 G   g 
X   l  Reg0.4  l    Cg 
 Gg    l 

0.5

 
 l  g 

0.5

zcoordinate along the tube direction, mGreek symbols







thermal conductivity, W m-1 K-1
viscosity, Pa·s
density, kg m-3
surface tension, N m-1
two-phase flow multiplier
Subscriptsggas-phase, vapor-phaseL,lliquid-phaseTP
two-phaseAcknowledgment
This study was sponsored by the project "Development of Non-fluorinated Energy-Saving
Refrigeration and Air Conditioning Systems" of New Energy and Industrial Technology
Development Organization, Japan. The authors thank Mr. T. Kamiaka of the University ofTokyo Graduate School for cooperation with calculating thermal property values.
References
Akasaka, R., Tanaka, K., Higashi, Y., 2010. Thermodynamic property modeling for-7-

2,3,3,3-tetrafluoropene (HFO-1234yf), Int. J. Refrigeration 33, 52–60.
Arakawa Y., S. Kim, T. Kamiaka, C. Dang, E. Hihara, 2010, Thermophysical property
measurement of HFO-1234yf +HFC-32 mixture. 2010 International Symposium on
Next-generation Air Conditioning and Refrigeration Technology, Feb. 17-19,Tokyo,Japan.
NS22.
Brown, J. S., Zilio, C., Cavallini, A., 2009. Estimations of the thermodynamic and transport
properties of R-1234yf using a cubic equation of state and group contribution methods. 3rd IIR
Conference on Thermophysical Properties and Transfer Processes of Refrigerants, June 23-26,
Boulder, Colorado, USA. IIR-127.
Chisholm, D., 1967. A theoretical basis for the Lockhart-Martinelli correlation for two-phaseflow. Int. J. Heat Mass Transfer 10, 1767–1778.
Gungor, K.E., Winterton, R.H.S., 1987. Simplified general correlation for saturated flow
boiling and comparisons of correlations. Chem. Eng. Research & Design 65, 148–156.
Hulse, R., Singh, R., Pham, H., 2009. Physical properties of HFO-1234yf. 3rd IIR Conferenceon Thermophysical Properties and Transfer Processes of Refrigerants, June 23-26, Boulder,
Colorado, USA. IIR-178.
Kandlikar, S.G., 1990. A general correlation for saturated two-phase flow boiling heattransfer inside horizontal and vertical tubes. ASME J. Heat Transfer 112, 219–228.
Lemmon, E.W., Huber, M.L., McLinden, M.O., 2007. NIST Thermodynamic and Transport
Properties of Refrigerants and Refrigerant Mixtures (REFPROP) Version 8.0. NIST.
Nicola, G. D., Palanara, F., Santori, G., 2010. Saturated pressure measurements of
2,3,3,3-tetrafluoroprop-1-ene (HFO-1234yf). J. Chem. Eng. Data 55, 2010204.
Poling, B.E., Prausnitz, J.M., O’Connell, J.P., 2001. The Properties of Gases and Liquids, 5thed. McGraw-Hill, New York.
Saitoh, S., Daiguji, H., Hihara, E., 2005. Effect of tube diameter on boiling heat transfer of
R-134a in horizontal small-diameter tubes. Int. J. Heat Mass Transfer 48, 4973–4984.
Saitoh, S., Daiguji, H., Hihara, E., 2007. Correlation for boiling heat transfer of R-134a inhorizontal tubes including effect of tube diameter. Int. J. Heat Mass Transfer 50, 5215–5225.
Tanaka, K., Higashi, Y., 2010. Thermodynamic properties of HFO-1234yf(2,3,3,3-tetrafluoropropene). Int. J. Refrigeration 33, 474–479.
Stephan, K., Abdelsalam M., 1980. Heat-transfer correlations for natural convection boiling,
Int. J. Heat Mass Transfer 23, 73–87.
Yoshida, S., Mori, H., Hong, H., Matsunaga, T. 1994. Prediction of heat transfer coefficient
for refrigerants flowing in horizontal evaporator tubes. Trans. of the JAR 11(1), 67–78.

-8-Figure 1. Schematic of experimental system used to measure flow boiling heat transfer.

-9-

P

P: pressure tap

a=52.75 mm

T: thermo couple

b=106 mm

P
T

a

T

b

T

b

T

T

b

a

a

T

b

T

b

T

b

a

P

P

P
T

a

T

b

T

b

b

T

T

a

a

1760

Test section
T-typethermocouple

Junction

Tube(SUS304)
EnamelTeflonsheet

Side view

Thermocouple attachment
Figure 2. Schematic of 2 mm ID test tube.

- 10 -

T

b

T

b

T

b

a

10
-2 -1

-2 -1

hexp [kW m K ]8-2

G=200 kg m s , q= 6 kW m
-2 -1
-2
G=200 kg m s , q=12 kW m
-2 -1
-2
G=200 kg m s , q=24 kW m

6

4

2

0
0.0

0.2

0.4

0.6

0.8

1.0

x

Figure 3. Effect of heat flux on local heat transfer coefficient.

- 11 -

10
-2 -1

-2 -1

hexp [kW m K ]8-2

G=100 kg m s , q=12 kW m
-2 -1
-2
G=200 kg m s , q=12 kW m
-2 -1
-2
G=400 kg m s , q=12 kW m

6

4

2

0
0.0

0.2

0.4

0.6

0.8

1.0

x

Figure 4. Effect of mass flux on local heat transfer coefficient.

- 12 -

10
-2 -1

-2 -1

hexp [kW m K ]8-2

R-134a, 300 kg m s , 12 kW m
-2 -1
-2
HFO-1234yf, 300 kg m s , 12 kW m

6

4

2

0
0.0

0.2

0.4

0.6

0.8

1.0

x

Figure 5. Comparison of heat transfer coefficients between HFO-1234yf and R-134a at massflux of 300 kg m-2 s-1 and heat flux of 12 kW m-2.

- 13 -

100
7
65liquid(turbulent flow)
liquid(laminar flow)

43

hexp /hL

2

10
7
6
5
4
3
2

1
1

hexp /hL∝(1/X)

2

3

4

0.77

5 6 7 89

10

2

3

4

5 6 7 89

100

1/X

Figure 6. Boiling heat transfer coefficient as function of Lockhart-Martinelli parameter- 14 -

10

10

＋20％
+20％

8

-2 -1

hexp [kW m K ]-2 -1

hexp [kW m K ]

8

－20％

6

4

－20％
4

2

2

0

6

-2 -1

Mass flux :100～400 kg m s-2Heat flux : 6～24 kW m

-2 -1

Mass flux :100～400 kg m s-2Heat flux : 6～24 kW m

0
0

2

4

6

8

10

0

2

4

6

hcal [kW m K ](a) Correlation of Saitoh et al.
10
＋20％

＋20％
8

－20％

-2 -1

6

hexp [kW m K ]-2 -1

hexp [kW m K ]

8

4

2

2

4

6

8

－20％

4

2

-2 -1

Mass flux :100～400 kg m s-2Heat flux : 6～24 kW m
0

6

-2 -1

Mass flux :100～400 kg m s-2Heat flux : 6～24 kW m
0

10

0

2

4

-2 -1

8

10

hcal [kW m K ](c) Correlation of Kandlikar(d) Correlation of Gungor-Winterton

8

8
-2

-1

-2

G=200 kg m s , q=12 kW m (measured)
Correlation of Saitoh et al.
Correlation of Yoshida
Correlation of Kandlikar
Correlation of Gungor-Winterton

6

2

2

h [kW/m K]

6

6
-2 -1

hcal [kW m K ]h [kW/m K]

10(b) Correlation of Yoshida

10

0

8

-2 -1

hcal [kW m K ]

-2 -1

4

4

-2

2

0
0.0

2

0.2

0.4

0.6

0.8

(e) Predicted and measured values forflux 200kg m-2 s-1 and heat flux 12kW m-2
x

0
0.0

1.0

-1

-2

G=300 kg m s , q=12 kW m (mesured)
Correlation of Saitoh et al.
Correlation of Yoshida
Correlation of Kandlikar
Correlation of Gungor-Winterton
0.2

0.4

0.6

0.8

1.0

(f) Predicted and measured values for massflux 300kg m-2 s-1 and heat flux 12kW m-2
x

Figure 7. Experimental flow boiling heat transfer coefficient hexp vs. calculated hcal for
HFO-1234yf, and comparison of four predicted and measured data.
- 15 -

80][kg m-2 s-1] [kW m-2

Pressure drop [kPa]60

exp.-100-6
cal. -100-6
exp.-100-12
cal. -100-12
exp.-200-6
cal. -200-6
exp.-200-12
cal. -200-12
exp.-200-24
cal. -200-24
exp.-400-12
cal. -400-12
exp.-400-24
cal. -400-24

40

20

0
0.0

0.5

1.0

1.5

2.0

Tube length [m]Figure 8. Comparison of pressure drops between measured values and calculated by
Lockhart-Martinelli correlation.

- 16 -

Table 1 Specification of test tube and measurement intervals
Material oftest tube
Stainless steel(SUS304)

ID oftube(mm)

OD oftube(mm)

Length oftube(mm)

T.C.

OD of
T.C.(mm)Interval of
T.C.(mm)Interval ofpressure(mm)

2

3

1760

T-type

0.1

106

450

Table 2 Experimental conditions
Refrigerant

Inlet temperature[°C]

Quality[1]

Heat flux[kW m-2]

Mass flux[kg m-2 s-1]

HFO-1234yf

15

0.1–1.0

6–24

100–400

Table 3 Saturation properties of HFO-1234yf and R-134a at temperature of 288.15 K

l

g

ilg

l

l

g

cPl



[kg m-3][kg m-3][kJ kg-1][W m-1 K-1]

[μPa s][μPa s][kJ kg-1 K-1]

[N m-1]

HFO-1234yf

1077.3

26.3

156.5

0.0724

177.7

11.80

1.337

0.0077

R-134a1243.423.76

186.590.0854220.6611.29

1.38690.0094

Table 4 Correlations for flow boiling heat transfer
Saitoh et al. ;1  We  , S  1 1  0.4  Re 10   , Re  Re F

hTP  Fhl  Shpool , F  1  1 X 

1.05

TP

 l  G 1  x  D   cpl l 

 
 ,
l
D
  l 
0.8

Rel  1000, hl  0.023
hpool  207

l  qd b 
d b  lTl 0.745

 g   l 

0.581

Prl

0.533

hTP  Fhl  Shpool , F  1  2 1 X 

0.88

 g   l 

0.581

Prl

0.533

4.36lD

 Bo 10   X , Bo  q Gh
4

Rel  1000, hl 



2

, d b  0.51  g  l  g  

0.5

Kandlikar ;0.5C 1  x   g 
hTP  hl C1Co C2  25 Frlo  5  C3 BoC4 FK  , Co  
  


 x   l 
for Co  0.65 : C1  1.136, C2  0.9, C3  667.2, C4  0.7, C5  0.3

Co  0.65 : C1  0.6683, C2  0.2, C3  1058, C4  0.7, C5  0.3
C 5  0 for vertical tubes, and for horizontal with Frl  0.04

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0.5

lv

0.4

0.8

1.25

l

0.5



 l  G 1  x  D   cpl l 

 
 ,
l
D
  l 

0.745

Rel  1000, hl 

0.5
, S  1 1  0.9  Retp 104 
0.8

  qd 
hpool  207 l  b 
d b  lTl 

TP

0.4



2
, d b  0.51  g  l  g  

Yoshida;Rel  1000, hl  0.023

4 1.40.4

g

4.36lD

Gungor and Winterton ;0.41

 l  g 
If the tube is horizontal and the Froude number Frl is less x 
hTP  Ehl , E  1  3000 Bo 0.86  1.12 1 x 0.75

than 0.05 then E should be multiplied by the factor
E 2  Frl 

0.1 2 Frl 

Table 5 Mean deviation and accuracy of three correlations for flow boiling heat transfer
coefficient
Correlation

Mean deviation(%)

Accuracy defined as
fraction of data within±20% error(%)

Saitoh et al.

9.2

92.8

Yoshida

20.4

51.9

Kandlikar

18.7

51.9

Gungor−Winterton

15.5

66.9

Mean deviation 

1
n



n
1

hexp  hcal
hexp

 100% , the number of data is 181.

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