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CO2 and R410A: Two-phase flow visualizations and flow boiling measurements at medium (0.50) reduced pressure
Applied Thermal Engineering 49 (2012) 2e8
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Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
CO2 and R410A: Two-phase flow visualizations and flow boiling measurements at medium (0.50) reduced pressure R. Mastrullo a, A.W. Mauro a, *, J.R. Thome b, G.P. Vanoli c a
DETEC, Università degli Studi di Napoli FEDERICO II, p.le Tecchio, 80, 80125, Napoli, Italy Laboratory of Heat and Mass Transfer (LTCM), Swiss Federal Institute of Technology, Lausanne (EPFL), Lausanne CH-1015, Switzerland c DING, University of Sannio, Palazzo ex INPS – Piazza Roma, 21 – 82100 Benevento, Italy b
a r t i c l e i n f o
a b s t r a c t
Article history: Received 4 November 2010 Accepted 11 October 2011 Available online 20 October 2011
The conversion to more environmentally friendly refrigerants is of prime importance for refrigeration and air-conditioning industries. A first option for the substitution of the fluids which use was banned could be natural refrigerants, such as CO2, or HFC mixtures with low global warming and zero ozone depletion impacts. To determine the heat transfer characteristics during flow boiling of CO2 a remarkable amount of literature on flow boiling of CO2 has been published, especially during the last decade. The experiments on this fluid pointed out that, at medium reduced pressure, the heat transfer coefficients trends versus vapor quality are quite different than those commonly encountered with other refrigerants at ordinary temperatures and several interpretations for this were proposed. In this work tests for CO2 and R410A are carried out to study the behaviour of the local heat transfer during evaporation at medium reduced pressure (0.50). The tests were run on a smooth, horizontal, stainless steel tube with an inner diameter equal to 6.00 mm, varying the mass flux between 200 and 350 kg m2 s1 and the heat flux between 5.0 and 20.0 kW m2 over the whole range of vapor qualities. It was verified by experiments that for R410A at a medium reduced pressure the peripheral local heat transfer coefficient trends, as the two-phase flow structures and flow regime transitions, are similar to those for CO2, at the same reduced pressure. Ó 2011 Elsevier Ltd. All rights reserved.
Keywords: Flow boiling CO2 R410A Reduced pressure Heat transfer
1. Introduction Due to environmental issues such as ozone depletion and global warming, the refrigeration, air-conditioning and heat pump industries have started the conversion to environmentally friendly refrigerants. Depending on the size of the systems, the operating and ambient conditions, different options are available among natural refrigerants and HFC mixtures. In the last decade the interest of scientific and industrial research has been focused on ammonia, carbon dioxide (R744), hydrocarbons and other HFC mixtures, such as R404A, R407C, R410A [1e3]. The application of HFC mixtures has been recently introduced also in cascade air-to-water heat pumps for sanitary water production, increasing the working temperature during evaporation at the high temperature side. More recently the use of CO2, in saturated conditions and as secondary fluid, has been considered in a wider range of temperature, due to low pumping energy required.
* Corresponding author. Tel.: þ39 081762198; fax: þ39 0812309364. E-mail address: [email protected] (A.W. Mauro). 1359-4311/$ e see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.applthermaleng.2011.10.021
In such systems the refrigerant side heat transfer resistance is not negligible for the sizing of evaporators with the accuracy currently required and one important aspect is the knowledge of two-phase characteristics during flow boiling, as confirmed by the large literature production on this theme (see, for example [4,5]). Among other fluids, flow boiling of carbon dioxide has been intensively investigated. In effect, this fluid evaporates at much higher reduced pressures than other refrigerants (at the same evaporation temperatures). Hence, the thermodynamic and transport properties are remarkably different than those of conventional fluids. In particular, the surface tension and the liquid viscosity are much lower, causing a large increase in nucleate contribution to evaporation, while decreasing pressure gradients. As pointed out by Bredesen et al. [6], using carbon dioxide the heat transfer efficiency at refrigerant side could be considerably enhanced by increasing heat flux without punishment in pressure loss. However, the CO2 heat transfer coefficient trends are completely different than those of other fluids at ordinary temperatures, as the works in literature show. This study refers to flow boiling in a smooth tube with 6.00 mm inner diameter and the objective is to report experiments for two
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
fluids: CO2 and R410A, showing two-phase flow pattern transitions and local heat transfer coefficients at medium reduced pressures (almost 0.50). In particular, in order to study the changes in the heat transfer coefficient (HTC) trends versus the vapor quality, varying the mass flux and the heat flux, an analysis of the heat transfer along the periphery of the tube and two-phase flow pattern visualizations are carried out, highlighting the influence of nucleate and convective contributions on heat transfer and the analogies in local heat transfer trends between CO2 and R410A. 2. State of the art The review work by Thome and Ribatski [7] shows a collection of works before 2005 concerning flow boiling inside tubes. Referring to macro channels it is possible to find studies on smooth and enhanced tubes, with mass fluxes ranging from 85 to 650 kg m2 s1, heat fluxes from 3 to 58 kW m2 and saturation temperatures from 25 C to 20 C. The authors pointed out the main flow boiling characteristics for CO2 (R744) in tubes, like the higher heat transfer coefficients with respect to the commonly used refrigerants. They showed that the heat transfer strongly depends on heat flux confirming a large nucleate boiling contribution; the heat transfer is substantially independent of mass flux. Moreover the trends of the HTCs (heat transfer coefficients) versus vapor quality for R744 have contrasting trends, dependently on the operating conditions. About the latter aspect, it is well confirmed by experiments that the trends for R744 HTCs versus vapor quality (at small vapor qualities) are increasing at low evaporating temperatures (reduced pressures lower than 0.33). After a threshold value of pressure, dependent also on the diameter and the mass flux, the trends of the HTCs are decreasing with the vapor quality. For this behaviour, a work by Yun et al. [8] first proposed an interesting interpretation consistent with experiments and based on a flow pattern transition. The authors pointed out that the higher reduced pressure for R744 causes a higher liquid hold-up at the early stage of evaporation and the liquid level is enough to establish an intermittent flow regime (in their work, for mass fluxes greater than 170 kg m2 s1 at 10 C with an inner tube diameter equal to 6.00 mm). Moreover, they argued that during evaporation the transition to annular flow determines an unstable and asymmetric annular flow due to low values of a modified Froude number, in the form presented by Kefer et al. [9]. The unstable flow and the enhanced liquid droplet entrainment, due to a relatively thicker liquid film at the top of the tube, could cause the occurrence of dry-out patches and, hence, a boiling crisis at low vapor qualities. This idea was not quantified in their work and their interpretation apparently is not consistent with the results of their analysis. In effect, the transition value from annular flow to the dry-out, they suggested (0.35e0.40), is greater than the vapor quality at which the decline of HTCs starts, (0.15e0.20), in the operating conditions of their study. Other works deal with this issue. In particular, the papers by Cheng et al. [10,11] defined a new model for R744 flow pattern transitions. In the first work [10] Cheng et al. adapted the original flow pattern based model for local heat transfer coefficients by Wojtan et al. [12] to R744. In particular, starting from the experimental results of independent databases and observing the falldown and the sharp-drop of HTCs with vapor quality, they fit the coefficients of the formulas presented in [12] for the transition from intermittent to annular flow and from annular flow to dry-out. Then, the authors proposed a new flow boiling model, developing a correlation for the nucleate boiling contribution in a form similar to Cooper’s one [13]. The formula accounts for the suppression of nucleative contribution to the boiling due to the reduction of the
3
temperature difference, and the corresponding superheating, at the wall. In the formula there is a factor that returns a value equal to 1 for the intermittent region and a value related to the mean liquid thickness after the transition to annular flow. In particular, the liquid film thickness was calculated with the expression proposed by El Hajal et al. [14]. Moreover, they did the same for the dry-out region. In the updated model [11] Cheng et al. proposed further modifications, but the basic ideas for the flow transitions and the suppression of nucleative contribution were confirmed. The models proposed in [10] and [11] do not account for the transition to the dry-out at an early stage of evaporation, as supposed by Yun et al. [8]. Nevertheless, slightly anticipating the transition from the intermittent to the annular flow regime and increasing the nucleate boiling contribution to the heat transfer, the trends of the models follow quite well the HTC experiments. Although the good predictive results, the models in [10] and [11] were not founded on a phenomenological basis, since no experimental evidence to justify the HTC changes in trends was yet available. In a more recent work [15] Mastrullo et al. carried out an analysis of the heat transfer characteristics of R744, analyzing the distribution of the local heat transfer coefficients along the periphery of the tube varying the operating conditions. They pointed out that, for the cases in which a decline of the HTCs was noted, the heat transfer coefficients on the periphery are almost constant except for the top of the tube, where a strong variation was encountered. Increasing the vapor quality at low mass flux (200 kg m2 s1 at 3 C in a 6.00 mm internal diameter tube) the decline of the top HTC leads to values lower than those at the bottom; instead at higher mass fluxes all the values are similar. The authors argued that at low vapor qualities a stratified-wavy or slug-flow regime is reasonably established, according to the interpretation proposed by Yun et al. [8] and by Cheng et al. [10]; a transition to annular flow is expected only at high mass fluxes (higher than 200 kg m2 s1, in that work). However, differently than Yun et al. [8], they did not suppose an early transition to dryout and in the slug region they found that HTCs increased differently to what results from the models by Cheng et al. [10] and [11], where the HTCs are supposed constant. In the slug region they proposed an interpretation of the trends at the top side of the tube, starting from the conclusions of the work by Sun et al. [16]. In particular, Sun et al. proposed a model dividing the flow in slug and film regions, where the heat transfer coefficient was averaged on the residence times of the slugs and the liquid films at the top of the tube. As a first step Sun et al. carried out an analysis of the hydrodynamics of the slug flow, using the model proposed by Coney [17], concluding that the film thickness in the slug region is almost constant, varying the vapor quality and the mass flux. Moreover, they found that the liquid at the top is moving slowly than that at the bottom in the average. The conclusions by Sun et al. [16] can fit also what happens along the evaporation of R744, according to Mastrullo et al. [15]. However in that work [15] no flow pattern visualizations was collected to support completely this conclusion. In this paper a study based on flow pattern visualizations and heat transfer coefficient measurements along the periphery of the tube varying the vapor quality, the mass flux and the heat flux at a medium reduced pressure is presented for R744. The experiments in similar conditions are repeated also for R410A to investigate analogies in heat transfer with R744 at similar reduced pressure. 3. Description of the experimental plant A schematic view of the experimental plant is shown in Fig. 1. The refrigerant loop consists of a magnetic gear pump, a pre-heater,
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an adiabatic and a diabatic test section, a shell-and-tube heat exchanger, a brazed plate heat exchanger and a tube-in-tube subcooler. The magnetic gear pump drives the fluid coming from the liquid reservoir. The refrigerant mass flux can be modified varying the electric motor speed by an inverter. The refrigerant, in subcooled conditions, passes first through the pre-heater where heat is supplied to the fluid by four fibreglass heating tapes (each one has a nominal power of 830 W at 240 V (AC)); changing the voltage it is possible to modify the thermal power and to obtain the desired vapor quality at the diabatic test section inlet. After the pre-heater, the fluid flows first through the adiabatic test section and then through the diabatic test section. At the exit of the diabatic test section a smooth glass tube 10.0 cm long, with the same inner diameter of the test tube allows to see the two-phase flow patterns and a high frequency of frames camera is installed to record the flow regimes in each experimental condition. The diabatic test section is a smooth, horizontal, circular, stainless steel (type 304) tube with an inner radius of 3.00 0.05 mm, an outer radius of 4.00 0.05 mm and a length of 780.0 0.5 mm at 20 C and 1.0 bar. The thermal power is provided to the fluid by Joule effect by a feed current device. After the evaporation, the refrigerant condenses in a shell-andtube heat exchanger and in a brazed plate heat exchanger. Before returning to the pump, the refrigerant is sub-cooled in a tube-intube heat exchanger. The coolant is an auxiliary fluid (aqua-glycol mixture) contained in a storage tank of 200 dm3. It can be chilled down to 30 C by a R404A auxiliary refrigerating plant and it is circulated by a magnetic gear pump connected to an inverter. By adjusting the refrigerant charge, the aqua-glycol mixture inlet temperature and mass flow rate, it is possible to modify and hold constant the refrigerant evaporating pressure in the test section. To avoid heat gains, heavy insulation was provided by an elastomeric insulator (k ¼ 0.035 W m1 K1 at 0.0 C) for the two test sections, shell-and-tube heat exchanger, the tube-in-tube sub-cooler, the liquid reservoir, the tubes and the tube fittings; by a 32 mm layer of cellular insulator (k ¼ 0.040 W m1 K1 at 40 C) for the plate heat exchanger and by 5 cm layer of rock wool insulator (k ¼ 0.075 W m1 K1 at 300 C) for the pre-heater. Preliminary tests were conducted in single-phase conditions to verify the error
in the evaluation of the fluid temperature by energy balance and direct measurements. An agreement always better than 2% was found and the assumption of negligible heat losses is acceptable. A more detailed description of the test section, data reduction, preliminary tests to check the consistency of the assumptions for data reduction are available in Mastrullo et al. [15]. At the exit of the diabatic test section, a smooth glass tube, 100 mm long, with the same inner diameter as the test tube, provides visual access to record the two-phase flow regimes by a high speed digital video camera at the high pressures encountered in this study. 4. Measurement method and data reduction The local heat transfer coefficient is measured in the section M, far away 140.0 0.05 mm from the exit of the diabatic test section by the Newton equation:
hi ¼
q tw;i tsat
(1)
where tw,i [ C] is calculated from the measured outside wall temperature tw,o [ C] by applying the one-dimensional, radial, steady-state heat conduction equation for a hollow cylinder with a uniform heat generation; tw,o [ C] is measured with four four-wire Pt100 U resistance thermometers, with an uncertainty of 0.15 C, mounted on the top, the bottom, the left and the right sides of the tube in order to taking into account the liquid and vapor spatial distributions; the heat flux is calculated as follows:
q ¼
QDTS 2pri;DTS LDTS
(2)
The heating power, QDTS [W], provided to the fluid by Joule effect, is calculated measuring the voltage VDTS between the inlet and the exit of the tube by an electrical voltage transducer of 0e10 V of the calibration range with an uncertainty of 0.19% of reading þ0.01% of full scale; the electrical resistance of tube RDTS from the calibration certificate is 0.0244479 U 0.037 mU at 5.00 C and 0.026491 0.040 mU at 30.00 C.
Fig. 1. Schematic view of the experimental test facility.
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
The vapor quality is determined by the refrigerant properties calculation software RefProp [18] with the pressure (directly measured with a piezoelectric absolute pressure transducer with a measurement uncertainty of 0.1 bar) and the enthalpy determined by an energy balance from the inlet of the pre-heater to the section M. The refrigerant mass flux is measured with a Coriolis mass flow meter with an uncertainty of 0.05% of the read value. The measurement of the power supplied at the pre-heater is provided by a wattmeter characterized by the uncertainty of 0.2% of the measured value and the 0.02% of the full scale. The pressures are measured with piezoelectric absolute pressure transducer of 0e50 bar with a 0.1% of full scale uncertainty. The logging of signals from all the sensors is performed on all the channels for 100 s with 1.0 Hz acquisition frequency and these values for each channel are stored. If the deviation of each value from its average value is lower than a fixed quantity (0.5 C for the tw,o, 0.1 bar for the saturation pressure, 10 kg m2 s1 for the mass flux and 0.5 kW m2 for the heat flux) steady-state conditions are assumed. Then the local heat transfer coefficients are calculated for the 100 data acquisition set of values and a time-averaged value is considered as the local value in each position (htop, hbott, hleft, hright) at the measurement location (section M in Fig. 1). These heat transfer coefficients are the local peripheral heat transfer coefficients. From the local peripheral heat transfer coefficients it is possible to calculate, the cross sectional averaged values, hlocal, defined as follows:
hlocal ¼
htop þ hbott þ hleft þ hright 4
(3)
The uncertainty propagation in the calculations are evaluated by Moffat [19] analysis. For the tests presented in this work the accuracy for each local heat transfer coefficient is always better than 7% and the absolute accuracy for vapor qualities is always better than 0.02. 5. Experimental results and discussion 5.1. Two-phase flow visualizations Table 1 reports two-phase flow pictures for R744 and R410A at a reduced pressure (pr) almost equal to 0.50 (saturation temperatures equal to 5 C for R744 and 42 C for R410A) varying the refrigerant mass flux (G) within the range from 200 to 350 kg m2 s1 at a constant heat flux (q) almost equal to 10.0 kW m2. In the table only the most relevant pictures to capture the flow pattern transitions varying the mass flux and the vapor quality are reported. The flow pattern identifications were done also by movies, not reported here. For R744 the experimental results showed that at the lower mass flux (200 kg m2 s1) there is a transition from slug to stratified-wavy flow and, then, a partial dry-out of the wetted perimeter is encountered. At mass fluxes higher than 200 kg m2 s1, the sequence of transitions in flow patterns is: slug, an unstable coexistence of stratified-wavy and asymmetric annular flows, annular flow and, finally, the dry-out. The asymmetric annular flow, mentioned here, refers to an annular flow with high liquid hold-up and a liquidevapor interface with a persistent formation of waves and vortexes. Also a high entrainment of liquid droplets in the vapor core is observed. This flow regime can be supposed as a superimposition of the stratifiedwavy and the annular flows. This result confirms the observations
5
by Yun et al. [8]. Moreover it is noted that the dry-out occurs at very high vapor qualities (>0.9). For R410A the most important difference with respect to R744 is that the transition to dry-out is anticipated; moreover in the unstable flow the stratified-wavy flow seems to be slightly predominant. In conclusion, at the same reduced pressure the flow patterns and their transitions for these fluids are quite similar. 5.2. Comparison of two-phase flow visualizations to two-phase flow pattern maps About the comparison of experiments to the flow pattern maps, in Table 1 also the predictions by the Cheng et al. map [11] for R744 and by the Wojtan et al. map [20] for R410A are reported. It is possible to notice that the maps work quite well, also that for R410A [20] that originally was not fitted on so high reduced pressure data. However, minor differences are encountered; especially the influence of the mass flux is not well captured in the transition from stratified-wavy to annular flow regime. Moreover, it was noticed that the transition from slug to intermittent flow regime is under-estimated by the flow pattern map by Cheng et al. [11]. For example, in the test with a mass flux equal to 350 kg m2 s1 the transition was found at a vapor quality equal to 0.37 and not 0.18, as expected according to the map. On the contrary, for the R410A the map by Wojtan et al. [20] slightly overestimates the transition boundary from intermittent to annular flow, while predicted quite well the dry-out inception. The models considered for the flow pattern map transitions are based on dimensionless numbers, which correlate the main thermodynamic properties. It is possible to compare the main thermodynamic properties and their combinations at the same reduced pressure to see which group of properties is much consistent with experiments. As showed in Table 2, the main thermodynamic properties have almost the same values both for liquid and vapor phases at the same reduced pressure; the only remarkable variation is for the surface tension. Nevertheless, their combination in dimensionless numbers is more important. Hence, the main dimensionless numbers of the Cheng et al. method [11] are calculated and reported in Table 3. It is possible to notice that even if the liquid Weber (WeL) and Froude (FrL) numbers are quite different, due to the influence of the surface tension, the void fractions (3 ), the dimensionless liquid level (hLD) and the dimensionless wetted liquid angle (QSTRAT/2p) are almost the same for the two fluids, consistently with the experiments. Hence, the latter group of parameters is more important in the determination of the flow pattern transition and strongly related to reduced pressure. 5.3. Heat transfer coefficient measurements Further considerations about the importance of reduced pressure for heat transfer and two-phase flow transitions are possible considering the heat transfer coefficients. In Fig. 2 the measurements of the local heat transfer coefficient versus vapor quality at an almost constant mass flux equal to 350 kg m2 s1 varying the reduced pressure (from 0.12 to 0.52) are reported for R410A. It is possible to notice that for low reduced pressures the typical trends of HTCs are encountered. In effect, in the low vapor quality region (up to 0.20) the influence of the nucleate boiling contribution is encountered, while after the transition to annular flow the HTCs increase monotonically, as for pure convection, due to the mean fluid velocity increase during evaporation. At medium reduced pressures there is a major importance of heat transfer at low vapor
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Table 1 Flow pattern visualizations for R410A and CO2 for mass fluxes equal to 200 and 350 kg/m2 s, at a reduced pressure equal to 0.50 and different vapor qualities compared to flow pattern predictions according to [20] for R410A and [11] for R744. Fluid
G [kg m2 s1]
Vapor quality
R744
200
R744
Two-phase flow picture
Flow pattern interpretation
Flow pattern prediction
0.19
Slug
Transition slug/annular
200
0.85
Dry-out
Dry-out
R744
350
0.18
Slug
Transition slug/annular
R744
350
0.48
Stratified-wavy and asymmetric annular
Annular
R744
350
0.82
Annular
Dry-out
R410A
200
0.18
Slug
Transition slug/intermittent
R410A
200
0.68
Stratified-wavy and asymmetric annular
Annular
R410A
350
0.23
Slug
Intermittent
R410A
350
0.62
Asymmetric annular
Annular
R410A
350
0.88
Annular
Dry-out
qualities; after the transition to annular flow the heat transfer coefficients decrease. Hence, the trend of the local heat transfer coefficient for R410A at medium reduced pressure is the same that for R744 (for example, see the data presented in [15]). This behaviour is observed also for other mass fluxes (data not reported here). It confirms that the reduced pressure influences the change in trends of the local heat transfer coefficients also for R410A, like for R744. To better explore what happens at medium reduced pressures we can consider the experiments reported in Fig. 3. In particular, in Fig. 3a there are the experiments for R744 at a reduced pressure (pr)
Table 2 Thermodynamic properties for R744 (CO2) and R410A at saturation for a reduced pressure equal to 0.50.
rL rV mL mV s
3
[kg m ] [kg m3] [106 Pa s] [106 Pa s] [N m1]
R744
R410A
% variation (ref. to 744)
896 115 90.8 þ15.4 3.62e03
955 110 92.0 þ15.2 3.00e03
þ6.6 3.9 þ1.3 1.3 17.3
equal to 0.57; in Fig. 3b there are the experiments for R401A at pr ¼ 0.52, combining for both the cases the mass fluxes 200 and 350 kg m2 s1 to the heat fluxes equal to 5.0 and 20.0 kW m2. For R744 (Fig. 3a) it is confirmed, as expected, that the heat flux has a much higher influence than mass flux on the HTCs, at medium reduced pressures. In particular, for the experimental conditions reported in this work, an increase on the heat flux determines the increase of the HTCs in the whole vapor quality range. Instead, an increase of the mass flux determines an increase of the HTCs only in
Table 3 Main parameters for transition from stratified to annular flow, according to the method by [11], for a reduced pressure equal to 0.50 and a vapor quality equal to 0.30. R744 G WeL FrL WeL/FrL 3
QSTRAT/2p hLD
[kg m2 s1] [e] [e] [e] [e] [e] [e]
200 74 0.846 87 0.654 0.58 0.378
R410A 350 226 2.59 87 0.677 0.59 0.359
200 84 0.745 112 0.673 0.59 0.362
350 257 2.28 112 0.695 0.60 0.344
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
relative importance are remarkably different. In particular, the main difference is that for R410A the values at the top of the tube are higher than those at the bottom. Since the nucleation needs enough liquid to avoid its suppression this contribution is mainly encountered at the bottom of the tube, where a large liquid hold-up was registered during two-phase flow visualizations; conversely at the top of the tube the liquid film is much thinner than at the bottom. These effects are encountered in the whole vapor quality range (both in the slug/intermittent region and the (asymmetric) annular region). For these reasons the top peripheral HTC is influenced strongly by changes in flow regimes (in this work changes in mass flux from 200 to 350 kg m2 s1 at fixed vapor quality), while the heat transfer at the bottom is mainly controlled by the nucleate contribution to the evaporation. The latter contribution is remarkably more important for R744 than R410A, as can be seen comparing Fig. 4a to Fig. 4b.
R410A - G = 350 kg m-2 s-1, q = 10.0 kW m-2 8.00 7.00
5.00 4.00 3.00 tsat= -8.9 °C, pr=0.12 tsat= +5.1 °C, pr=0.19 tsat= +30.1 °C, pr=0.39
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality Fig. 2. Effect on heat transfer of the variation from low (0.12) to medium (0.52) reduced pressure for R410A: local heat transfer coefficients versus vapor quality at a mass flux almost equal to 350 kg m2 s1 and a heat flux equal to 10.0 kW m2.
the medium-high vapor quality range (almost after 0.35) and before the dry-out inception. The influence of mass flux and/or heat flux variations on dry-out inception is confirmed. For R410A (Fig. 3b), unlike R744, the mass flux has a much higher influence than heat flux on the HTCs, at medium reduced pressures. In particular, for the experimental conditions reported in this work, an increase of the heat flux causes an increase of the HTCs at vapor qualities lower than 0.45 and G ¼ 200 kg m2 s1; the same increase of the heat flux at G ¼ 350 kg m2 s1 determines an increase of HTCs almost in the whole range of vapor qualities (before the dry-out inception). An influence of the heat flux on the dry-out inception was encountered only at G ¼ 350 kg m2 s1: higher the heat flux lower the dry-out inception vapor quality, as expected. Focusing the attention on the trends of the HTCs versus vapor quality, almost the same behaviour is encountered when comparing Fig. 3a (R744) to Fig. 3b (R410A), except for R744 G ¼ 350 kg m2 s1 and q ¼ 5.0 kW m2. In effect, during evaporation the HTCs have a decrease up to a vapor quality equal to 0.40e0.45; then they are almost constant up to the dry-out inception. While at G ¼ 350 kg m2 s1 the HTCs have at first a decrease, then an increase, before the dry-out inception. In order to carry out a deeper analysis of these trends, the local peripheral HTCs are measured in a fixed cross section of the tube at four locations (top, bottom, right and left). Fig. 4 reports the measurements of the peripheral HTCs only along the vertical direction (in effect, Fig. 3 reports the spatial average of all the four local peripheral HTCs, but the strong variations occurs along vertical direction reported in Fig. 4). In particular, Fig. 4a refers to R744, pr ¼ 0.57 and G ¼ 200 kg m2 s1 varying the heat flux from 5.0e20.0 kW m2; Fig. 4b refers to R410A, pr ¼ 0.52 and G ¼ 200 kg m2 s1 varying the heat flux from 5.0e20.0 kW m2. Fig. 4a shows that the local peripheral HTCs at the bottom is almost constant or slightly increasing during evaporation, for R744; at the top the HTC is decreasing up to a vapor quality equal to 0.40, then it is almost constant. The heat transfer at the top of the tube is clearly influenced by the changes in two-phase flow patterns from slug flow to annular flow. Moreover the value at the bottom is higher than that at the top for high vapor qualities. Fig. 4b shows that for R410A the local peripheral HTCs have almost the same trends than for R744. The trends at the top and the bottom of the tube are similar, but the absolute values and their
a
G = 200 kg m-2 s-1, q = 5 kW m-2 G = 350 kg m-2 s-1, q = 5 kW m-2 G = 200 kg m-2 s-1, q = 20 kW m-2 G = 350 kg m-2 s-1, q = 20 kW m-2
R744 - tsat = 7.0 °C, pr = 0.57
24.0
20.0
-2 -1 hlocal [kW m K ]
0.00 0.0
tsat= +42.1 °C, pr=0.52
16.0
12.0
8.0
4.0
0.0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality
b
G = 200 kg m-2 s-1, q = 5 kW m-2 G = 350 kg m-2 s-1, q = 5 kW m-2 G = 200 kg m-2 s-1, q = 20 kW m-2 G = 350 kg m-2 s-1, q = 20 kW m-2
R410A - tsat = 41.7 °C, pr = 0.52
8.00
6.00
-2 -1
1.00
hlocal [kW m K ]
hlocal [kW m-2 K-1]
6.00
2.00
7
4.00
2.00
0.00 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality Fig. 3. Local heat transfer coefficients versus vapor quality at a reduced pressure almost equal to 0.50, for two mass fluxes (200 and 350 kg m2 s1) and two heat fluxes (5.0 and 20.0 kW m2), for: (2a) R744 and (2b) R410A.
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a
q = 5 kW m-2, htop
R744 - tsat = 7.0 °C, pr = 0.57,
q = 5 kW m-2, hbott
G = 200 kg m-2 s-1
q = 20 kW m-2, htop q = 20 kW m-2, hbott
28.0
-2 -1
h peripheral[kW m K ]
24.0 20.0 16.0 12.0 8.0 4.0 0.0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
experiments in conditions similar to CO2 are repeated for R410A to investigate analogies in flow patterns and heat transfer with R744. The flow pattern visualizations confirmed the presence of a high liquid hold-up and hence the existence of an asymmetric annular flow, that has a large influence on the heat transfer for both the fluids. Moreover, it was verified that, at the same reduced pressure, the flow pattern transitions and the two-phase flow structures are the same both for R744 and R410A. Concerning the heat transfer the analysis of the local peripheral heat transfer coefficients showed that the trends versus the vapor quality are almost the same for the two fluids, with strong difference in their absolute values and their relative weights on the spatial average value. In particular, the heat transfer coefficients at the bottom of the tube are mainly influenced by nucleative contribution due to the high liquid hold-up, and hence it has the predominant effect for CO2, remarking the strong dependence on the heat flux and the almost independence on the mass flux for this fluid. On the contrary, for R410A the convective contribution is the most important, and so the heat transfer is mainly influenced by the heat transfer at the top of the tube, strongly related to the flow pattern transitions.
Vapor Quality
b
References q = 5 kW m-2, htop
R410A - tsat = 41.7 °C, pr = 0.52,
q = 5 kW m-2, hbott
G = 200 kg m-2 s-1
q = 20 kW m-2, htop q = 20 kW m-2, hbott
12.0
-2 -1
hperipheral [kW m K ]
10.0
8.0
6.0
4.0
2.0
0.0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality Fig. 4. Local peripheral heat transfer coefficients at the top and the bottom of the tube versus vapor quality, at a reduced pressure almost equal to 0.50 and a mass flux almost equal to 200 kg m2 s1, varying the heat flux from 5.0 to 20.0 kW m2 for: (3a) R744 and (3b) R410A.
In conclusion, each local peripheral HTC trend (Fig. 4) has the same behaviour for R744 and R410A, but their spatial average (Fig. 3) is different due to their different relative weights.
6. Conclusions This paper refers to flow boiling in a smooth tube with 6.00 mm inner diameter and reports experiments for two fluids: CO2 and R410A, at medium reduced pressures (almost 0.50). The study was based on flow pattern visualizations and heat transfer coefficient measurements along the periphery of the tube varying the vapor quality, the mass flux and the heat flux. The
[1] G. Lorentzen, The use of natural refrigerants: a complete solution to the CFC/ HCFC predicament, Int. J. Refrigeration 18 (1995) 190e197. [2] G. Lorentzen, Revival of carbon dioxide as a refrigerant, Int. J. Refrigeration 17 (1994) 292e301. [3] A. Cavallini, Working fluids for mechanical refrigeration invited paper presented at the 19th international congress of refrigeration, The Hague, August 1995, Int. J. Refrigeration 19 (1996) 485e496. [4] J.R. Thome, Boiling of new refrigerants: a state-of-the-art review, Int. J. Refrigeration 19 (1996) 435e457. [5] J.R. Thome, L. Cheng, G. Ribatski, L.F. Vales, Flow boiling of ammonia and hydrocarbons: a state-of-the-art review, Int. J. Refrigeration 31 (2008) 603e620. [6] A.M. Bredesen, A.H. Hafner, J. Pettersen, P. Neksa, K. Aflekt, Heat Transfer and Pressure Drop for in Tube Evaporation of CO2. IIF-IIR-Commission B1, with E1 & E2, College Park, 1997, pp. 35e49. [7] J.R. Thome, G. Ribatski, State-of-the-art of two-phase flow and flow boiling heat transfer and pressure drop of CO2 in macro- and micro-channels, Int. J. Refrigeration 28 (2005) 1149e1168. [8] R. Yun, Y. Kim, M.S. Kim, Youndon Choi, Boiling heat transfer and dry-out phenomenon of CO2 in a horizontal smooth tube, Int. J. Heat Mass Transfer 46 (2003) 2353e2361. [9] V. Kefer, W. Köhler, W. Kastner, Critical heat flux and post-CHF heat transfer in horizontal and inclined evaporator tubes, Int. J. Multiphase Flow 15 (1989) 385e392. [10] L. Cheng, G. Ribatski, L. Wojtan, J.R. Thome, New flow boiling heat transfer model and flow pattern map for carbon dioxide evaporating inside horizontal tubes, Int. J. Heat Mass Transfer 49 (2006) 4082e4094. [11] L. Cheng, G. Ribatski, J.R. Thome, New prediction methods for CO2 evaporation inside tubes: part II e an updated general flow boiling heat transfer model based on flow patterns, Int. J. Heat Mass Transfer 51 (2008) 125e135. [12] L. Wojtan, T. Ursenbacker, J.R. Thome, Investigation of flow boiling in horizontal tubes: part II e development of a new heat transfer model for stratified-wavy, dry-out and mist flow regimes, Int. J. Heat Mass Transfer 48 (2005) 2970e2985. [13] M.K. Cooper, Saturated nucleate pool boiling: a simple correlation, , In: First U.K. National Heat Transfer Conference, vol. 2 (1984) 785e793. [14] J.R. Thome, J. El Hajal, Flow boiling heat transfer to carbon dioxide: general prediction method, Int. J. Refrigeration 28 (2005) 1149e1168. [15] R. Mastrullo, A.W. Mauro, A. Rosato, G.P. Vanoli, Carbon dioxide local heat transfer coefficients during flow boiling in a horizontal smooth tube, Int. J. Heat Mass Transfer 52 (2009) 4184e4194. [16] G. Sun, G.F. Hewitt, V.V. Wadekar, A heat transfer for slug flow in a horizontal tube, Int. J. Heat Mass Transfer 47 (2004) 2807e2816. [17] M.W.E. Coney, The analysis of a mechanism of liquid replenishment and draining in horizontal two-phase flow, Int. J. Multiphase Flow 1 (1974). [18] E.W. Lemmon, M.O. Mc Linden, M.L. Huber, NIST Standard Reference Database 23, Version 8.0. REFPROP, 2007. [19] R.J. Moffat, Describing the uncertainties in experimental results, Exp. Thermal Fluid Sci. 1 (1988) 3e17. [20] L. Wojtan, T. Ursenbacher, J.R. Thome, Investigation of flow boiling in horizontal tubes: part I e a new diabatic two-phase flow pattern map, Int. J. Heat Mass Transfer 48 (2005) 2955e2969.
Contents lists available at SciVerse ScienceDirect
Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
CO2 and R410A: Two-phase flow visualizations and flow boiling measurements at medium (0.50) reduced pressure R. Mastrullo a, A.W. Mauro a, *, J.R. Thome b, G.P. Vanoli c a
DETEC, Università degli Studi di Napoli FEDERICO II, p.le Tecchio, 80, 80125, Napoli, Italy Laboratory of Heat and Mass Transfer (LTCM), Swiss Federal Institute of Technology, Lausanne (EPFL), Lausanne CH-1015, Switzerland c DING, University of Sannio, Palazzo ex INPS – Piazza Roma, 21 – 82100 Benevento, Italy b
a r t i c l e i n f o
a b s t r a c t
Article history: Received 4 November 2010 Accepted 11 October 2011 Available online 20 October 2011
The conversion to more environmentally friendly refrigerants is of prime importance for refrigeration and air-conditioning industries. A first option for the substitution of the fluids which use was banned could be natural refrigerants, such as CO2, or HFC mixtures with low global warming and zero ozone depletion impacts. To determine the heat transfer characteristics during flow boiling of CO2 a remarkable amount of literature on flow boiling of CO2 has been published, especially during the last decade. The experiments on this fluid pointed out that, at medium reduced pressure, the heat transfer coefficients trends versus vapor quality are quite different than those commonly encountered with other refrigerants at ordinary temperatures and several interpretations for this were proposed. In this work tests for CO2 and R410A are carried out to study the behaviour of the local heat transfer during evaporation at medium reduced pressure (0.50). The tests were run on a smooth, horizontal, stainless steel tube with an inner diameter equal to 6.00 mm, varying the mass flux between 200 and 350 kg m2 s1 and the heat flux between 5.0 and 20.0 kW m2 over the whole range of vapor qualities. It was verified by experiments that for R410A at a medium reduced pressure the peripheral local heat transfer coefficient trends, as the two-phase flow structures and flow regime transitions, are similar to those for CO2, at the same reduced pressure. Ó 2011 Elsevier Ltd. All rights reserved.
Keywords: Flow boiling CO2 R410A Reduced pressure Heat transfer
1. Introduction Due to environmental issues such as ozone depletion and global warming, the refrigeration, air-conditioning and heat pump industries have started the conversion to environmentally friendly refrigerants. Depending on the size of the systems, the operating and ambient conditions, different options are available among natural refrigerants and HFC mixtures. In the last decade the interest of scientific and industrial research has been focused on ammonia, carbon dioxide (R744), hydrocarbons and other HFC mixtures, such as R404A, R407C, R410A [1e3]. The application of HFC mixtures has been recently introduced also in cascade air-to-water heat pumps for sanitary water production, increasing the working temperature during evaporation at the high temperature side. More recently the use of CO2, in saturated conditions and as secondary fluid, has been considered in a wider range of temperature, due to low pumping energy required.
* Corresponding author. Tel.: þ39 081762198; fax: þ39 0812309364. E-mail address: [email protected] (A.W. Mauro). 1359-4311/$ e see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.applthermaleng.2011.10.021
In such systems the refrigerant side heat transfer resistance is not negligible for the sizing of evaporators with the accuracy currently required and one important aspect is the knowledge of two-phase characteristics during flow boiling, as confirmed by the large literature production on this theme (see, for example [4,5]). Among other fluids, flow boiling of carbon dioxide has been intensively investigated. In effect, this fluid evaporates at much higher reduced pressures than other refrigerants (at the same evaporation temperatures). Hence, the thermodynamic and transport properties are remarkably different than those of conventional fluids. In particular, the surface tension and the liquid viscosity are much lower, causing a large increase in nucleate contribution to evaporation, while decreasing pressure gradients. As pointed out by Bredesen et al. [6], using carbon dioxide the heat transfer efficiency at refrigerant side could be considerably enhanced by increasing heat flux without punishment in pressure loss. However, the CO2 heat transfer coefficient trends are completely different than those of other fluids at ordinary temperatures, as the works in literature show. This study refers to flow boiling in a smooth tube with 6.00 mm inner diameter and the objective is to report experiments for two
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
fluids: CO2 and R410A, showing two-phase flow pattern transitions and local heat transfer coefficients at medium reduced pressures (almost 0.50). In particular, in order to study the changes in the heat transfer coefficient (HTC) trends versus the vapor quality, varying the mass flux and the heat flux, an analysis of the heat transfer along the periphery of the tube and two-phase flow pattern visualizations are carried out, highlighting the influence of nucleate and convective contributions on heat transfer and the analogies in local heat transfer trends between CO2 and R410A. 2. State of the art The review work by Thome and Ribatski [7] shows a collection of works before 2005 concerning flow boiling inside tubes. Referring to macro channels it is possible to find studies on smooth and enhanced tubes, with mass fluxes ranging from 85 to 650 kg m2 s1, heat fluxes from 3 to 58 kW m2 and saturation temperatures from 25 C to 20 C. The authors pointed out the main flow boiling characteristics for CO2 (R744) in tubes, like the higher heat transfer coefficients with respect to the commonly used refrigerants. They showed that the heat transfer strongly depends on heat flux confirming a large nucleate boiling contribution; the heat transfer is substantially independent of mass flux. Moreover the trends of the HTCs (heat transfer coefficients) versus vapor quality for R744 have contrasting trends, dependently on the operating conditions. About the latter aspect, it is well confirmed by experiments that the trends for R744 HTCs versus vapor quality (at small vapor qualities) are increasing at low evaporating temperatures (reduced pressures lower than 0.33). After a threshold value of pressure, dependent also on the diameter and the mass flux, the trends of the HTCs are decreasing with the vapor quality. For this behaviour, a work by Yun et al. [8] first proposed an interesting interpretation consistent with experiments and based on a flow pattern transition. The authors pointed out that the higher reduced pressure for R744 causes a higher liquid hold-up at the early stage of evaporation and the liquid level is enough to establish an intermittent flow regime (in their work, for mass fluxes greater than 170 kg m2 s1 at 10 C with an inner tube diameter equal to 6.00 mm). Moreover, they argued that during evaporation the transition to annular flow determines an unstable and asymmetric annular flow due to low values of a modified Froude number, in the form presented by Kefer et al. [9]. The unstable flow and the enhanced liquid droplet entrainment, due to a relatively thicker liquid film at the top of the tube, could cause the occurrence of dry-out patches and, hence, a boiling crisis at low vapor qualities. This idea was not quantified in their work and their interpretation apparently is not consistent with the results of their analysis. In effect, the transition value from annular flow to the dry-out, they suggested (0.35e0.40), is greater than the vapor quality at which the decline of HTCs starts, (0.15e0.20), in the operating conditions of their study. Other works deal with this issue. In particular, the papers by Cheng et al. [10,11] defined a new model for R744 flow pattern transitions. In the first work [10] Cheng et al. adapted the original flow pattern based model for local heat transfer coefficients by Wojtan et al. [12] to R744. In particular, starting from the experimental results of independent databases and observing the falldown and the sharp-drop of HTCs with vapor quality, they fit the coefficients of the formulas presented in [12] for the transition from intermittent to annular flow and from annular flow to dry-out. Then, the authors proposed a new flow boiling model, developing a correlation for the nucleate boiling contribution in a form similar to Cooper’s one [13]. The formula accounts for the suppression of nucleative contribution to the boiling due to the reduction of the
3
temperature difference, and the corresponding superheating, at the wall. In the formula there is a factor that returns a value equal to 1 for the intermittent region and a value related to the mean liquid thickness after the transition to annular flow. In particular, the liquid film thickness was calculated with the expression proposed by El Hajal et al. [14]. Moreover, they did the same for the dry-out region. In the updated model [11] Cheng et al. proposed further modifications, but the basic ideas for the flow transitions and the suppression of nucleative contribution were confirmed. The models proposed in [10] and [11] do not account for the transition to the dry-out at an early stage of evaporation, as supposed by Yun et al. [8]. Nevertheless, slightly anticipating the transition from the intermittent to the annular flow regime and increasing the nucleate boiling contribution to the heat transfer, the trends of the models follow quite well the HTC experiments. Although the good predictive results, the models in [10] and [11] were not founded on a phenomenological basis, since no experimental evidence to justify the HTC changes in trends was yet available. In a more recent work [15] Mastrullo et al. carried out an analysis of the heat transfer characteristics of R744, analyzing the distribution of the local heat transfer coefficients along the periphery of the tube varying the operating conditions. They pointed out that, for the cases in which a decline of the HTCs was noted, the heat transfer coefficients on the periphery are almost constant except for the top of the tube, where a strong variation was encountered. Increasing the vapor quality at low mass flux (200 kg m2 s1 at 3 C in a 6.00 mm internal diameter tube) the decline of the top HTC leads to values lower than those at the bottom; instead at higher mass fluxes all the values are similar. The authors argued that at low vapor qualities a stratified-wavy or slug-flow regime is reasonably established, according to the interpretation proposed by Yun et al. [8] and by Cheng et al. [10]; a transition to annular flow is expected only at high mass fluxes (higher than 200 kg m2 s1, in that work). However, differently than Yun et al. [8], they did not suppose an early transition to dryout and in the slug region they found that HTCs increased differently to what results from the models by Cheng et al. [10] and [11], where the HTCs are supposed constant. In the slug region they proposed an interpretation of the trends at the top side of the tube, starting from the conclusions of the work by Sun et al. [16]. In particular, Sun et al. proposed a model dividing the flow in slug and film regions, where the heat transfer coefficient was averaged on the residence times of the slugs and the liquid films at the top of the tube. As a first step Sun et al. carried out an analysis of the hydrodynamics of the slug flow, using the model proposed by Coney [17], concluding that the film thickness in the slug region is almost constant, varying the vapor quality and the mass flux. Moreover, they found that the liquid at the top is moving slowly than that at the bottom in the average. The conclusions by Sun et al. [16] can fit also what happens along the evaporation of R744, according to Mastrullo et al. [15]. However in that work [15] no flow pattern visualizations was collected to support completely this conclusion. In this paper a study based on flow pattern visualizations and heat transfer coefficient measurements along the periphery of the tube varying the vapor quality, the mass flux and the heat flux at a medium reduced pressure is presented for R744. The experiments in similar conditions are repeated also for R410A to investigate analogies in heat transfer with R744 at similar reduced pressure. 3. Description of the experimental plant A schematic view of the experimental plant is shown in Fig. 1. The refrigerant loop consists of a magnetic gear pump, a pre-heater,
4
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
an adiabatic and a diabatic test section, a shell-and-tube heat exchanger, a brazed plate heat exchanger and a tube-in-tube subcooler. The magnetic gear pump drives the fluid coming from the liquid reservoir. The refrigerant mass flux can be modified varying the electric motor speed by an inverter. The refrigerant, in subcooled conditions, passes first through the pre-heater where heat is supplied to the fluid by four fibreglass heating tapes (each one has a nominal power of 830 W at 240 V (AC)); changing the voltage it is possible to modify the thermal power and to obtain the desired vapor quality at the diabatic test section inlet. After the pre-heater, the fluid flows first through the adiabatic test section and then through the diabatic test section. At the exit of the diabatic test section a smooth glass tube 10.0 cm long, with the same inner diameter of the test tube allows to see the two-phase flow patterns and a high frequency of frames camera is installed to record the flow regimes in each experimental condition. The diabatic test section is a smooth, horizontal, circular, stainless steel (type 304) tube with an inner radius of 3.00 0.05 mm, an outer radius of 4.00 0.05 mm and a length of 780.0 0.5 mm at 20 C and 1.0 bar. The thermal power is provided to the fluid by Joule effect by a feed current device. After the evaporation, the refrigerant condenses in a shell-andtube heat exchanger and in a brazed plate heat exchanger. Before returning to the pump, the refrigerant is sub-cooled in a tube-intube heat exchanger. The coolant is an auxiliary fluid (aqua-glycol mixture) contained in a storage tank of 200 dm3. It can be chilled down to 30 C by a R404A auxiliary refrigerating plant and it is circulated by a magnetic gear pump connected to an inverter. By adjusting the refrigerant charge, the aqua-glycol mixture inlet temperature and mass flow rate, it is possible to modify and hold constant the refrigerant evaporating pressure in the test section. To avoid heat gains, heavy insulation was provided by an elastomeric insulator (k ¼ 0.035 W m1 K1 at 0.0 C) for the two test sections, shell-and-tube heat exchanger, the tube-in-tube sub-cooler, the liquid reservoir, the tubes and the tube fittings; by a 32 mm layer of cellular insulator (k ¼ 0.040 W m1 K1 at 40 C) for the plate heat exchanger and by 5 cm layer of rock wool insulator (k ¼ 0.075 W m1 K1 at 300 C) for the pre-heater. Preliminary tests were conducted in single-phase conditions to verify the error
in the evaluation of the fluid temperature by energy balance and direct measurements. An agreement always better than 2% was found and the assumption of negligible heat losses is acceptable. A more detailed description of the test section, data reduction, preliminary tests to check the consistency of the assumptions for data reduction are available in Mastrullo et al. [15]. At the exit of the diabatic test section, a smooth glass tube, 100 mm long, with the same inner diameter as the test tube, provides visual access to record the two-phase flow regimes by a high speed digital video camera at the high pressures encountered in this study. 4. Measurement method and data reduction The local heat transfer coefficient is measured in the section M, far away 140.0 0.05 mm from the exit of the diabatic test section by the Newton equation:
hi ¼
q tw;i tsat
(1)
where tw,i [ C] is calculated from the measured outside wall temperature tw,o [ C] by applying the one-dimensional, radial, steady-state heat conduction equation for a hollow cylinder with a uniform heat generation; tw,o [ C] is measured with four four-wire Pt100 U resistance thermometers, with an uncertainty of 0.15 C, mounted on the top, the bottom, the left and the right sides of the tube in order to taking into account the liquid and vapor spatial distributions; the heat flux is calculated as follows:
q ¼
QDTS 2pri;DTS LDTS
(2)
The heating power, QDTS [W], provided to the fluid by Joule effect, is calculated measuring the voltage VDTS between the inlet and the exit of the tube by an electrical voltage transducer of 0e10 V of the calibration range with an uncertainty of 0.19% of reading þ0.01% of full scale; the electrical resistance of tube RDTS from the calibration certificate is 0.0244479 U 0.037 mU at 5.00 C and 0.026491 0.040 mU at 30.00 C.
Fig. 1. Schematic view of the experimental test facility.
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
The vapor quality is determined by the refrigerant properties calculation software RefProp [18] with the pressure (directly measured with a piezoelectric absolute pressure transducer with a measurement uncertainty of 0.1 bar) and the enthalpy determined by an energy balance from the inlet of the pre-heater to the section M. The refrigerant mass flux is measured with a Coriolis mass flow meter with an uncertainty of 0.05% of the read value. The measurement of the power supplied at the pre-heater is provided by a wattmeter characterized by the uncertainty of 0.2% of the measured value and the 0.02% of the full scale. The pressures are measured with piezoelectric absolute pressure transducer of 0e50 bar with a 0.1% of full scale uncertainty. The logging of signals from all the sensors is performed on all the channels for 100 s with 1.0 Hz acquisition frequency and these values for each channel are stored. If the deviation of each value from its average value is lower than a fixed quantity (0.5 C for the tw,o, 0.1 bar for the saturation pressure, 10 kg m2 s1 for the mass flux and 0.5 kW m2 for the heat flux) steady-state conditions are assumed. Then the local heat transfer coefficients are calculated for the 100 data acquisition set of values and a time-averaged value is considered as the local value in each position (htop, hbott, hleft, hright) at the measurement location (section M in Fig. 1). These heat transfer coefficients are the local peripheral heat transfer coefficients. From the local peripheral heat transfer coefficients it is possible to calculate, the cross sectional averaged values, hlocal, defined as follows:
hlocal ¼
htop þ hbott þ hleft þ hright 4
(3)
The uncertainty propagation in the calculations are evaluated by Moffat [19] analysis. For the tests presented in this work the accuracy for each local heat transfer coefficient is always better than 7% and the absolute accuracy for vapor qualities is always better than 0.02. 5. Experimental results and discussion 5.1. Two-phase flow visualizations Table 1 reports two-phase flow pictures for R744 and R410A at a reduced pressure (pr) almost equal to 0.50 (saturation temperatures equal to 5 C for R744 and 42 C for R410A) varying the refrigerant mass flux (G) within the range from 200 to 350 kg m2 s1 at a constant heat flux (q) almost equal to 10.0 kW m2. In the table only the most relevant pictures to capture the flow pattern transitions varying the mass flux and the vapor quality are reported. The flow pattern identifications were done also by movies, not reported here. For R744 the experimental results showed that at the lower mass flux (200 kg m2 s1) there is a transition from slug to stratified-wavy flow and, then, a partial dry-out of the wetted perimeter is encountered. At mass fluxes higher than 200 kg m2 s1, the sequence of transitions in flow patterns is: slug, an unstable coexistence of stratified-wavy and asymmetric annular flows, annular flow and, finally, the dry-out. The asymmetric annular flow, mentioned here, refers to an annular flow with high liquid hold-up and a liquidevapor interface with a persistent formation of waves and vortexes. Also a high entrainment of liquid droplets in the vapor core is observed. This flow regime can be supposed as a superimposition of the stratifiedwavy and the annular flows. This result confirms the observations
5
by Yun et al. [8]. Moreover it is noted that the dry-out occurs at very high vapor qualities (>0.9). For R410A the most important difference with respect to R744 is that the transition to dry-out is anticipated; moreover in the unstable flow the stratified-wavy flow seems to be slightly predominant. In conclusion, at the same reduced pressure the flow patterns and their transitions for these fluids are quite similar. 5.2. Comparison of two-phase flow visualizations to two-phase flow pattern maps About the comparison of experiments to the flow pattern maps, in Table 1 also the predictions by the Cheng et al. map [11] for R744 and by the Wojtan et al. map [20] for R410A are reported. It is possible to notice that the maps work quite well, also that for R410A [20] that originally was not fitted on so high reduced pressure data. However, minor differences are encountered; especially the influence of the mass flux is not well captured in the transition from stratified-wavy to annular flow regime. Moreover, it was noticed that the transition from slug to intermittent flow regime is under-estimated by the flow pattern map by Cheng et al. [11]. For example, in the test with a mass flux equal to 350 kg m2 s1 the transition was found at a vapor quality equal to 0.37 and not 0.18, as expected according to the map. On the contrary, for the R410A the map by Wojtan et al. [20] slightly overestimates the transition boundary from intermittent to annular flow, while predicted quite well the dry-out inception. The models considered for the flow pattern map transitions are based on dimensionless numbers, which correlate the main thermodynamic properties. It is possible to compare the main thermodynamic properties and their combinations at the same reduced pressure to see which group of properties is much consistent with experiments. As showed in Table 2, the main thermodynamic properties have almost the same values both for liquid and vapor phases at the same reduced pressure; the only remarkable variation is for the surface tension. Nevertheless, their combination in dimensionless numbers is more important. Hence, the main dimensionless numbers of the Cheng et al. method [11] are calculated and reported in Table 3. It is possible to notice that even if the liquid Weber (WeL) and Froude (FrL) numbers are quite different, due to the influence of the surface tension, the void fractions (3 ), the dimensionless liquid level (hLD) and the dimensionless wetted liquid angle (QSTRAT/2p) are almost the same for the two fluids, consistently with the experiments. Hence, the latter group of parameters is more important in the determination of the flow pattern transition and strongly related to reduced pressure. 5.3. Heat transfer coefficient measurements Further considerations about the importance of reduced pressure for heat transfer and two-phase flow transitions are possible considering the heat transfer coefficients. In Fig. 2 the measurements of the local heat transfer coefficient versus vapor quality at an almost constant mass flux equal to 350 kg m2 s1 varying the reduced pressure (from 0.12 to 0.52) are reported for R410A. It is possible to notice that for low reduced pressures the typical trends of HTCs are encountered. In effect, in the low vapor quality region (up to 0.20) the influence of the nucleate boiling contribution is encountered, while after the transition to annular flow the HTCs increase monotonically, as for pure convection, due to the mean fluid velocity increase during evaporation. At medium reduced pressures there is a major importance of heat transfer at low vapor
6
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
Table 1 Flow pattern visualizations for R410A and CO2 for mass fluxes equal to 200 and 350 kg/m2 s, at a reduced pressure equal to 0.50 and different vapor qualities compared to flow pattern predictions according to [20] for R410A and [11] for R744. Fluid
G [kg m2 s1]
Vapor quality
R744
200
R744
Two-phase flow picture
Flow pattern interpretation
Flow pattern prediction
0.19
Slug
Transition slug/annular
200
0.85
Dry-out
Dry-out
R744
350
0.18
Slug
Transition slug/annular
R744
350
0.48
Stratified-wavy and asymmetric annular
Annular
R744
350
0.82
Annular
Dry-out
R410A
200
0.18
Slug
Transition slug/intermittent
R410A
200
0.68
Stratified-wavy and asymmetric annular
Annular
R410A
350
0.23
Slug
Intermittent
R410A
350
0.62
Asymmetric annular
Annular
R410A
350
0.88
Annular
Dry-out
qualities; after the transition to annular flow the heat transfer coefficients decrease. Hence, the trend of the local heat transfer coefficient for R410A at medium reduced pressure is the same that for R744 (for example, see the data presented in [15]). This behaviour is observed also for other mass fluxes (data not reported here). It confirms that the reduced pressure influences the change in trends of the local heat transfer coefficients also for R410A, like for R744. To better explore what happens at medium reduced pressures we can consider the experiments reported in Fig. 3. In particular, in Fig. 3a there are the experiments for R744 at a reduced pressure (pr)
Table 2 Thermodynamic properties for R744 (CO2) and R410A at saturation for a reduced pressure equal to 0.50.
rL rV mL mV s
3
[kg m ] [kg m3] [106 Pa s] [106 Pa s] [N m1]
R744
R410A
% variation (ref. to 744)
896 115 90.8 þ15.4 3.62e03
955 110 92.0 þ15.2 3.00e03
þ6.6 3.9 þ1.3 1.3 17.3
equal to 0.57; in Fig. 3b there are the experiments for R401A at pr ¼ 0.52, combining for both the cases the mass fluxes 200 and 350 kg m2 s1 to the heat fluxes equal to 5.0 and 20.0 kW m2. For R744 (Fig. 3a) it is confirmed, as expected, that the heat flux has a much higher influence than mass flux on the HTCs, at medium reduced pressures. In particular, for the experimental conditions reported in this work, an increase on the heat flux determines the increase of the HTCs in the whole vapor quality range. Instead, an increase of the mass flux determines an increase of the HTCs only in
Table 3 Main parameters for transition from stratified to annular flow, according to the method by [11], for a reduced pressure equal to 0.50 and a vapor quality equal to 0.30. R744 G WeL FrL WeL/FrL 3
QSTRAT/2p hLD
[kg m2 s1] [e] [e] [e] [e] [e] [e]
200 74 0.846 87 0.654 0.58 0.378
R410A 350 226 2.59 87 0.677 0.59 0.359
200 84 0.745 112 0.673 0.59 0.362
350 257 2.28 112 0.695 0.60 0.344
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
relative importance are remarkably different. In particular, the main difference is that for R410A the values at the top of the tube are higher than those at the bottom. Since the nucleation needs enough liquid to avoid its suppression this contribution is mainly encountered at the bottom of the tube, where a large liquid hold-up was registered during two-phase flow visualizations; conversely at the top of the tube the liquid film is much thinner than at the bottom. These effects are encountered in the whole vapor quality range (both in the slug/intermittent region and the (asymmetric) annular region). For these reasons the top peripheral HTC is influenced strongly by changes in flow regimes (in this work changes in mass flux from 200 to 350 kg m2 s1 at fixed vapor quality), while the heat transfer at the bottom is mainly controlled by the nucleate contribution to the evaporation. The latter contribution is remarkably more important for R744 than R410A, as can be seen comparing Fig. 4a to Fig. 4b.
R410A - G = 350 kg m-2 s-1, q = 10.0 kW m-2 8.00 7.00
5.00 4.00 3.00 tsat= -8.9 °C, pr=0.12 tsat= +5.1 °C, pr=0.19 tsat= +30.1 °C, pr=0.39
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality Fig. 2. Effect on heat transfer of the variation from low (0.12) to medium (0.52) reduced pressure for R410A: local heat transfer coefficients versus vapor quality at a mass flux almost equal to 350 kg m2 s1 and a heat flux equal to 10.0 kW m2.
the medium-high vapor quality range (almost after 0.35) and before the dry-out inception. The influence of mass flux and/or heat flux variations on dry-out inception is confirmed. For R410A (Fig. 3b), unlike R744, the mass flux has a much higher influence than heat flux on the HTCs, at medium reduced pressures. In particular, for the experimental conditions reported in this work, an increase of the heat flux causes an increase of the HTCs at vapor qualities lower than 0.45 and G ¼ 200 kg m2 s1; the same increase of the heat flux at G ¼ 350 kg m2 s1 determines an increase of HTCs almost in the whole range of vapor qualities (before the dry-out inception). An influence of the heat flux on the dry-out inception was encountered only at G ¼ 350 kg m2 s1: higher the heat flux lower the dry-out inception vapor quality, as expected. Focusing the attention on the trends of the HTCs versus vapor quality, almost the same behaviour is encountered when comparing Fig. 3a (R744) to Fig. 3b (R410A), except for R744 G ¼ 350 kg m2 s1 and q ¼ 5.0 kW m2. In effect, during evaporation the HTCs have a decrease up to a vapor quality equal to 0.40e0.45; then they are almost constant up to the dry-out inception. While at G ¼ 350 kg m2 s1 the HTCs have at first a decrease, then an increase, before the dry-out inception. In order to carry out a deeper analysis of these trends, the local peripheral HTCs are measured in a fixed cross section of the tube at four locations (top, bottom, right and left). Fig. 4 reports the measurements of the peripheral HTCs only along the vertical direction (in effect, Fig. 3 reports the spatial average of all the four local peripheral HTCs, but the strong variations occurs along vertical direction reported in Fig. 4). In particular, Fig. 4a refers to R744, pr ¼ 0.57 and G ¼ 200 kg m2 s1 varying the heat flux from 5.0e20.0 kW m2; Fig. 4b refers to R410A, pr ¼ 0.52 and G ¼ 200 kg m2 s1 varying the heat flux from 5.0e20.0 kW m2. Fig. 4a shows that the local peripheral HTCs at the bottom is almost constant or slightly increasing during evaporation, for R744; at the top the HTC is decreasing up to a vapor quality equal to 0.40, then it is almost constant. The heat transfer at the top of the tube is clearly influenced by the changes in two-phase flow patterns from slug flow to annular flow. Moreover the value at the bottom is higher than that at the top for high vapor qualities. Fig. 4b shows that for R410A the local peripheral HTCs have almost the same trends than for R744. The trends at the top and the bottom of the tube are similar, but the absolute values and their
a
G = 200 kg m-2 s-1, q = 5 kW m-2 G = 350 kg m-2 s-1, q = 5 kW m-2 G = 200 kg m-2 s-1, q = 20 kW m-2 G = 350 kg m-2 s-1, q = 20 kW m-2
R744 - tsat = 7.0 °C, pr = 0.57
24.0
20.0
-2 -1 hlocal [kW m K ]
0.00 0.0
tsat= +42.1 °C, pr=0.52
16.0
12.0
8.0
4.0
0.0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality
b
G = 200 kg m-2 s-1, q = 5 kW m-2 G = 350 kg m-2 s-1, q = 5 kW m-2 G = 200 kg m-2 s-1, q = 20 kW m-2 G = 350 kg m-2 s-1, q = 20 kW m-2
R410A - tsat = 41.7 °C, pr = 0.52
8.00
6.00
-2 -1
1.00
hlocal [kW m K ]
hlocal [kW m-2 K-1]
6.00
2.00
7
4.00
2.00
0.00 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality Fig. 3. Local heat transfer coefficients versus vapor quality at a reduced pressure almost equal to 0.50, for two mass fluxes (200 and 350 kg m2 s1) and two heat fluxes (5.0 and 20.0 kW m2), for: (2a) R744 and (2b) R410A.
8
R. Mastrullo et al. / Applied Thermal Engineering 49 (2012) 2e8
a
q = 5 kW m-2, htop
R744 - tsat = 7.0 °C, pr = 0.57,
q = 5 kW m-2, hbott
G = 200 kg m-2 s-1
q = 20 kW m-2, htop q = 20 kW m-2, hbott
28.0
-2 -1
h peripheral[kW m K ]
24.0 20.0 16.0 12.0 8.0 4.0 0.0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
experiments in conditions similar to CO2 are repeated for R410A to investigate analogies in flow patterns and heat transfer with R744. The flow pattern visualizations confirmed the presence of a high liquid hold-up and hence the existence of an asymmetric annular flow, that has a large influence on the heat transfer for both the fluids. Moreover, it was verified that, at the same reduced pressure, the flow pattern transitions and the two-phase flow structures are the same both for R744 and R410A. Concerning the heat transfer the analysis of the local peripheral heat transfer coefficients showed that the trends versus the vapor quality are almost the same for the two fluids, with strong difference in their absolute values and their relative weights on the spatial average value. In particular, the heat transfer coefficients at the bottom of the tube are mainly influenced by nucleative contribution due to the high liquid hold-up, and hence it has the predominant effect for CO2, remarking the strong dependence on the heat flux and the almost independence on the mass flux for this fluid. On the contrary, for R410A the convective contribution is the most important, and so the heat transfer is mainly influenced by the heat transfer at the top of the tube, strongly related to the flow pattern transitions.
Vapor Quality
b
References q = 5 kW m-2, htop
R410A - tsat = 41.7 °C, pr = 0.52,
q = 5 kW m-2, hbott
G = 200 kg m-2 s-1
q = 20 kW m-2, htop q = 20 kW m-2, hbott
12.0
-2 -1
hperipheral [kW m K ]
10.0
8.0
6.0
4.0
2.0
0.0 0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0.9
1.0
Vapor Quality Fig. 4. Local peripheral heat transfer coefficients at the top and the bottom of the tube versus vapor quality, at a reduced pressure almost equal to 0.50 and a mass flux almost equal to 200 kg m2 s1, varying the heat flux from 5.0 to 20.0 kW m2 for: (3a) R744 and (3b) R410A.
In conclusion, each local peripheral HTC trend (Fig. 4) has the same behaviour for R744 and R410A, but their spatial average (Fig. 3) is different due to their different relative weights.
6. Conclusions This paper refers to flow boiling in a smooth tube with 6.00 mm inner diameter and reports experiments for two fluids: CO2 and R410A, at medium reduced pressures (almost 0.50). The study was based on flow pattern visualizations and heat transfer coefficient measurements along the periphery of the tube varying the vapor quality, the mass flux and the heat flux. The
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