Thermal oscillations during flow boiling of hydrocarbon refrigerants in a microchannels array heat sink

Thermal oscillations during flow boiling of hydrocarbon refrigerants in a microchannels array heat sink

Applied Thermal Engineering 157 (2019) 113725 Contents lists available at ScienceDirect Applied Thermal Engineering journal homepage: www.elsevier.c...

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Applied Thermal Engineering 157 (2019) 113725

Contents lists available at ScienceDirect

Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng

Research Paper

Thermal oscillations during flow boiling of hydrocarbon refrigerants in a microchannels array heat sink

T



Cristian A. Cháveza, , Nelson O. Moragaa, Gherhardt Ribatskib a b

Department of Mechanical Engineering, University of La Serena (ULS), La Serena, Chile Heat Transfer Research Group, Department of Mechanical Engineering, Escola de Engenharia de São Carlos (EESC), University of São Paulo (USP), São Carlos, SP, Brazil

H I GH L IG H T S

oscillations for R600a, R290 and R1270 during flow boiling in a microchannels array. • Thermal effects of the signal behaviors for temperature and pressure transducers. • Parametric of signal amplitudes based on the equilibrium between inertial and boiling forces. • Analysis • Refrigerant R1270 presents lower thermal oscillations for the same operational conditions than R600a and R290.

A R T I C LE I N FO

A B S T R A C T

Keywords: Thermal instabilities Flow boiling Multi-microchannels Reverse flow Natural fluids

This paper presents an experimental investigation on thermal oscillations during flow boiling of the hydrocarbons R600a, R290 and R1270 in a microchannels array heat sink. The test section is composed of fifty microchannels manufactured in a copper block with each channel having a rectangular cross-section of 123 × 494 μm2 and 15 mm in length. Experimental data were obtained for mass velocities of 165–823 kg/m2 s, heat fluxes up to 400 kW/m2, liquid subcooling at the inlet of the test section of 5, 10 and 15 °C and saturation temperatures of 21 and 25 °C. Fluctuations of the fluid temperature at the test section outlet and of the pressure at the heat sink inlet are presented and analyzed. The results reveal that the thermal oscillation amplitude increases with decreasing saturation temperature and increasing heat flux and liquid subcooling at the inlet of the test section. Propylene was found to present the lowest thermal oscillations. The analysis of thermal oscillations allows to identify potential conditions that allows to maximize the efficiency of the heat sink for microcooling applications.

1. Introduction and literature review In the last two decades, the development of electronic processors has been characterized by the increase in consumed power and dissipated heat as a result of the increment in the number of transistors inside these devices [1]. In addition, components of microelectromechanical systems (MEMS) such as some actuators as well as microreactors work under conditions of micrometric dimensions and high heat fluxes up to 1 kW/cm2 [2]. Such scenario implied on a drastic increase in the number of studies aiming to the development of devices and technologies capable of dissipating high heat fluxes in small spaces [3–17]. In this context, the technology of flow boiling in microchannels provides a suitable solution to such demands by combining reduced refrigerant inventory, high degree of compactness, high heat transfer coefficients and almost uniform fluid temperature along the heat ⁎

transfer process. However, some engineering challenges should be overcome before this technology become available in the market such as the presence of thermal instabilities and maldistribution effects [18–20]. The flow boiling instabilities are undesired phenomena which may cause a premature critical heat flux (CHF), high pressure drops, control and operational problems and mechanical vibrations of the system components [21,22]. These phenomena are generally segregated as static and dynamic instabilities. Tibiriçá et al. [20] describe the thermal instabilities in microchannels array heat sink as dynamic ones, particularly highlighting oscillations caused by compressible volumes, reverse flow and interaction of the flow in the plenum region. Reverse flow is associated to the microchannels connectivity and the abrupt bubble growth under confined conditions and their effects on the thermal instabilities are still an open issue. According to [23], the

Corresponding author. E-mail addresses: [email protected] (C.A. Chávez), [email protected] (N.O. Moraga), [email protected] (G. Ribatski).

https://doi.org/10.1016/j.applthermaleng.2019.113725 Received 26 December 2018; Received in revised form 9 April 2019; Accepted 3 May 2019 Available online 06 May 2019 1359-4311/ © 2019 Elsevier Ltd. All rights reserved.

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Nomenclature f G p Q̇ H q″ T W

2ϕ cont ch elec end env exp fluid fp f H in or 2 mo l lv M out or 3 v plms sat wall

friction factor [dimensionless] mass velocity [kg/m2 s] absolute pressure [kPa] heat [W] microchannel depth [m] heat flux [W/m2] temperature [°C] microchannel width [m]

Greek symbols Δp ζ ρ σ

differential pressure [kPa] ratio between microchannel and plenum areas [dimensionless] density [kg/m3] surface tension [N/m]

Subscripts 1ϕ

flow boiling condition contraction channel electric end environment expansion fluid footprint area frictional high heat sink inlet plenum momentum liquid-phase latent vaporization measured heat sink outlet plenum vapor-phase plenums saturated state wall

single-phase flow

three operational regions segregated according to the ratio of heat flux and mass velocity. They indicated that for a ratio (q/G) less than 0.95 kJ/kg the operational conditions correspond to a region of stable flow. Megahed [29] has verified the effect of mass velocity variation on thermal oscillations. For single microchannel, he observed an increment of the oscillation period with decreasing mass velocity, as it has been reported in the following Refs. [30–32]. Recently, Kuang et al. [33] noticed the same influence of mass velocity during flow boiling in an aluminum heat sink with four microchannels. Furthermore, they found that as the saturation temperature increased, the frequency of the thermal oscillations increased. In general, the frequency and amplitude of the oscillations increase with increasing heat flux. This behavior is due to the intensification of the boiling forces, [34]. The majority of the studies in literature concerning flow boiling evaporation in microchannels array heat sinks has been performed for parallel rectangular channels [35]. Generally, under some operating conditions, such a configuration provides a non-uniform distribution of the fluid, enhancing thermal instability as verified by [36]. In this context, Refs. [37–39] proposed configurations based on perpendicular cross-sectional channels with the purpose of achieving a uniform the fluid distribution, reducing the effects of confined bubble growth [32]. Configurations involving branched channels distributed according to fractals were also proposed by [40–43]. In addition to these studies to reduce the reverse flow solutions that involve the inclusion of microcolumns in the upstream region and restrictions at the entrance of the microchannel were implemented in [44,45]. An evaluation of the effect of these restrictions at the entrance of the microchannels has been described in [11],[36] and [46,47]. Solutions based on micro-holes promoting artificial nucleation sites along the channels were analyzed in [48,49]. Parallel microchannels containing a divergent cross section were proposed in [50–54]. This geometric feature restricts the return of the vapor phase to the input plenum. The effect on the thermal instabilities of microchannels orientation for horizontal, vertical upward, vertical downward and inclined channels was investigated in [16] and [55,56]. According to these authors, the lowest thermal instability effect, characterized by improved fluid distribution, no reverse flow presence and an increasing the heat transfer coefficient is observed for upward flow with the microchannel array heat sink positioned vertically. Recently, Ref. [57] compared the performance between parallel microchannel and an interconnected microchannels net. They observed

microchannels connectivity plays an important role on the reverse flow. In this case, the bubble nucleated under confined conditions grows and achieves the inlet plenum. Then, the two-phase back flow mixes with the inlet fluid, increasing the temperature of the fluid in the inlet plenum, reducing the saturation temperature and intensifying pressure fluctuations. Hetsroni et al. [4] investigated the process of bubble growth. They pointed out that the boiling forces are responsible for the flow in both directions along the microchannels configured in parallel. Besides, the authors observed small oscillations of temperature and pressure for single-phase flows. However, once verified the onset of nucleate boiling (ONB) the oscillations increased. Qu and Mudawar [24] and Lee and Mudawar [25] have studied the thermal instabilities effects through the analysis of the temperature and pressure oscillations in a multi-microchannel heat sink. Lee and Mudawar [25] measured the oscillations by including needle valves just upstream and downstream of the microchannels inlet and outlet, respectively. According to them, the amplitude of pressure oscillations was reduced by manipulating the needle valves, however, the instabilities due to reverse flow were not suppressed because they are inherent to the multi-microchannel and plenums configurations. In this regard, Kandlikar et al. [26] have described the effects of the bubble growth under confined conditions on critical heat flux and heat transfer coefficient during flow boiling inmicrochannels array heat sink. According to the authors, the non-controlled boiling forces combined with the microchannels geometry plus the plenum configuration and inlet effects destabilize the flow decreasing the heat sink performance. Kandlikar et al. [27] have proposed a model to predict the forces around the bubble due to the occurrence of thermal instabilities. According to them, the reverse flow presence can be evaluated through a balance between the inertial forces of the liquid and the boiling forces around a bubble. Subsequently, Lee et al. [11] modified the previous model by a microchannel with a divergent cross section and including the effects of a flow restriction at the channel inlet. In general, this modified model is based on an instability parameter given as a function of the mass velocity, heat flux, channel geometry and the fluid properties. A parameter of instability less than one was proposed as a criterion to define a stable flow condition. In the same context, Wang et al. [28] developed a map that characterizes stable and unstable operational conditions based on simultaneous visualization of flow boiling and measurement of heat flux and mass velocity. The authors defined 2

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lower oscillations of the wall temperature and fluid inlet temperature and pressure for the heat sink with a net of interconnected microchannels. Few studies are reported in literature focusing on the effects of fluid characteristics on thermal oscillations during flow boiling in microchannels heat sinks. Wang et al. [58] compared ethanol and FC-72, reporting wall temperature fluctuations with higher amplitude and lower wall temperature gradients in the channel axial direction for ethanol. In the study with FC-72 the wall temperature fluctuated chaotically and significant wall temperature gradients were noticed in the channel axial direction. According to the authors, these behaviors are associated to the lower surface tension of FC-72 and the complete wetting of the glass walls. Leão et al. [59] have performed a study on heat transfer and thermal oscillations of R407C and R134a. Their analyses of temperature at the outlet plenum signal by Fast Fourier Transform (FFT) revealed oscillations of lower amplitude and frequency for R407C compared to those of R134a. Currently, studies related to the fluid characteristic effects on thermal instabilities are required to enhance heat transfer and thermal efficiency with minimized pressure drops. In industry and academy, there is a tendency to look for refrigerants with excellent environmental properties and high efficiency [17]. For this reason, the chlorofluorocarbons (CFCs), hydrochlorofluorocarbons (HCFCs) and hydrofluorocarbons (HFCs) are being discarded as refrigerants due to their contribution to the ozone layer depletion and greenhouse effect. To

replace them, various alternatives were proposed and evaluated, as hydrofluoroolefins (HFOs) and the natural refrigerants (hydrocarbons, CO2 and ammonia). As pointed out by Thome [60], despite of some concerns about flammable refrigerants due to safety aspects, propane and isobutane are widely used as refrigerants in applications involving cascade refrigeration systems in ethylene production facilities and domestic refrigerators. In this context, it is important to highlight the almost negligible refrigerant charge contained in a microchannels heat sink developed for electronic cooling, corresponding to 0.15 g for the configuration estimated based on the dissipator dimensions using the homogeneous model to estimate the void fraction along the channels (see Chavez et al. [17] for details on this estimative). Despite of that, as pointed out by Chávez et al. [17], most of the studies on flow boiling in heat sinks based on multi-microchannels were performed for halocarbon refrigerants and water. Moreover, according to our best knowledge, the thermal oscillations performance of hydrocarbons evaporating across multi-microchannels heat sinks has not yet been evaluated. In this context, this study presents novel results concerning to characterizing thermal oscillations during flow boiling in a microchannel array heat sink for R600a, R290 and R1270. The R600a is a medium pressure refrigerant, while the R290 and R1270 are high pressure refrigerants. These natural fluids can be considered suitable in order to replace the CFCs, HCFCs and HFCs due to their null contribution to ozone layer depletion and low greenhouse effect. Focus was

Fig. 1. Experimental setup and test section details. 3

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using a VeccoWiko NT110 optical profiler at three positions of the channel side and the following arithmetic average results were obtained: arithmetical mean deviation Ra = 0.29 µm, root mean squared Rq = 0.35 µm and maximum height of the profile Rt = 3.23 µm. A Pyrex sheet is used to seal the test section on its upper part, as shown in Fig. 1. Four circular channels with an internal diameter of 3.5 mm were machined in the Pyrex sheet. Two of these channels are used to supply and drain the refrigerant from the test section. The additional pair of channels allows to connect an absolute and differential pressure transducer to the inlet and outlet plenums of the test section in order to measure the inlet pressure (p2) and the pressure drop (Δp) along the microchannels. Additionally, through these channels, thermocouple wires are introduced to register the fluid temperature at the inlet and outlet plenums (T2, T3). The needle Valve II is installed to prevent propagation of instabilities from the test section to the main circuit and vice versa. The ball Valve III is used to isolate the test section in case of carrying out maintenance procedures. Downstream the needle Valve II a plate-type heat exchanger is responsible to condense and sub cool the test fluid. In this heat exchanger, heat is dissipated to a water solution with 30% of ethylene glycol that circulates in a secondary circuit. The components of the apparatus and all its tubing are insulated with a layer of 20 mm of Armaflex elastomeric foam. The control and operation of the experimental setup is performed from a personal computer through a graphical interface programmed in LabView [61]. This interface communicated with an SCXI 1102 board plate installed in a National Instruments SCXI-1000 chassis which collects, processes and stores the electrical signal measurements of the instruments. The experimental data presented in this paper were obtained by gradually increasing the heat flux until a maximum and then decreasing its value down to zero. Therefore, in one experiment two sets data were measured. Each measurement corresponds to a 1000-point signal for which the Fast Fourier Transform (FFT) was calculated. It is important to highlight that experiments were repeated twice, with similar results obtained. The temperatures were measured using type K thermocouples with wire diameters of 150 µm. The absolute pressures were measured using piezoresistive transducers with a measurement interval of 0 to 4 bar for the experiments with R600a (low pressure) and with a measurement interval of 0 to 16 bar transducer used to experiments with R290 and R1270 (high pressure). The same test facility was used in the previous studies of Do Nascimento et al. [13], Leão et al. [16] and Chávez et al. [17], in which more detailed descriptions of the test apparatus are given.

given to analyze comparatively the effects of reduced pressure on the temperature fluctuations evaluated through a thermocouple located at the outlet plenum. The pressure oscillations at the inlet plenum were also evaluated through the signal provided by an absolute pressure transducer at the inlet of the test section. The experimental data cover heat fluxes based on the footprint area up to 400 kW/m2, mass velocities between 165 and 823 kg/m2 s and liquid subcooling at the inlet of the test section of 5, 10 and 15 °C and saturation temperatures of 21 and 25 °C. Results were carefully and parametrically analyzed considering the fluids effects on the fluctuations of temperature and pressure at the outlet and inlet plenum, respectively.

2. Experimental apparatus and procedure 2.1. Experimental setup and test section Fig. 1 presents a schematic of the refrigerant loop and details of the test section. In the refrigerant circuit, the fluid is driven by an oil-free gear micropump and the mass flow rate is measured by a Coriolis mass flow meter. The micropump is powered by a variable frequency drive that through the data acquisition system and based on the signal from the mass flow meters, controls the mass flow rate. The needle Valve I is installed in the circuit line, upstream the Pre-heater, to reduce the propagation of instability effects from the circuit to the test section and vice versa. A pre-heater made of a copper tube with a diameter of 12.7 mm is wrapped with tape-type electrical resistances. This device is installed upstream of the test section to heat up the refrigerant to the desired inlet subcooling temperature. A scheme of the experimental apparatus highlighting the test section components is depicted in Fig. 1. Furthermore, Fig. 2 depicts a detailed view of the microchannel array heat sinks. The fifty microchannels were made of a 28 × 25 × 4 mm3 copper block through a micro milling process using a tool with a 5 µm of precision. The microchannels had an average cross section of 123.3 × 494.2 µm2 and 15 mm in length. The separation fins between two consecutive channels was 176.9 µm. The base of the test section is composed by a Bakelite support block and the heating system is a 0.5 Ω electrical resistance made of a continuous Kanthal wire turned as serpentine. The electrical resistance is powered by a DC power source and electrically insulated from the test section by a mica layer at the bottom and a sapphire layer on top. The empty spaces between the sapphire and mica layers are filled with alumina. It is worth mention that the sapphire thermal conductivity is one order of magnitude higher than the mica thermal conductivity. This characteristic is associated to the fact that the main contribution of the heat flux is driven to the copper block containing the microchannels. Moreover, the relatively high conductivity of the sapphire wafer associated to its contact thermal resistance with the copper block helps to promote a uniform heat flux distribution on the heat sink base. The surface roughness was measured by Leão et al. [16]

2.2. Data reduction procedure A data reduction procedure was implemented to calculate from the measured variables the parameters that affect the behavior of the thermal oscillations. In this analysis, uniform mass velocity among the

Fig. 2. Heat sink details, (a) top view; (b) dimensions of microchannels. 4

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Fig. 4. Comparison of experimental and predicted Nusselt Number for singlephase flow.

Fig. 3. Comparison of experimental and predicted Fanning friction factors for single-phase flow.

1/2

Fforward ⎞ R=⎛ ⎝ Fback ⎠

microchannel and heat flux on the heat sink cooper surface contacting the fluid were assumed. Nonetheless, as indicated in a previous work of Chávez et al. [17], flow maldistribution, vapor back flows and flow instabilities were observed. However, the only feasible way to analyze the experimental results is assuming the aforementioned hypotheses because the local heat flux and mass velocity through each microchannel are not measurables. The mass velocity was calculated as the ratio between the mass flow given by the Coriolis type flowmeter and fifty times the cross-sectional ″ area of a microchannel. The heat flux referred to the footprint area qfp was calculated as the ratio between the estimated amount of heat reaching the microchannels Q̇ real given by Eq. (1) and the heated footprint area (15x15 mm2).

̇ = Qelect ̇ ̇ ̇ Qreal − Qplms − Qenv



si ∥s∥

Fforward =

G 2Ach ρL

(5)

and Fback is associated to the boiling force given by:

Fback =

̇ 1 ⎛ Qch ⎞⎟ ⎜ ρV ⎝ iLV ·2A ⎠

(6)

2.3. Validation of experimental data and uncertanity analyses The validation of experimental data and procedure were performed through the comparison between the single-phase flow pressure drop data and well-established predictive methods from literature [17,62,64,65]. For this purpose, the entrance and exit pressure drops due to the contraction and expansion at the inlet and outlet plenums, respectively, were subtracted from the total measured pressure drop ΔpM as follows:

(1)

Δp1ϕ = ΔpM - Δpj; ch

(7)

The entrance and exit pressure drops due to the contraction and expansion at the inlet and outlet plenums were estimated as proposed by Chalfi and Ghiaasiaan [64], respectively, according to the following equations:

Δpj; ch = Kj

G2 ; with j = \{ in,out\} 2ρj

(8)

where 2

ζ−1 ⎞ + 1 − ζ2 Kin = ⎜⎛ ⎟ ⎝ 1.08(ζ − 1) + 0.5371 ⎠

(2)

where sn is the normalized amplitude signal; si is the amplitude signal of temperature at the outlet plenum or pressure at the inlet plenum; ∥s∥ is the norm of amplitude signal calculated by following equation:

K out = 2ζ (ζ − 1)

∑ si2 i=1

(9) (10)

Fig. 3 presents a comparison between the Fanning friction factor based on the experimental pressure drop measurements and predictions through the method of Shah and London [65] for rectangular channels and laminar developing flow. Fig. 4 compares the experimental Nusselt Number and prediction values calculated through the model of Stephan and Preußer [62]. According to these figures, the experimental data and the prediction curves agree reasonably well, and it can be concluded that the experimental procedure is reasonable accurate. On the other hand, experiments were performed under similar conditions at the beginning and the end of each series of experiments. Thus, the

n

∥s∥ =

(4)

where Fforward is associated to the inertial force calculated as follows:

where Q̇ elect is the power supplied by the DC power source to the electrical resistance, estimated as the product between the voltage and electrical current supplied to the Kanthal wire, which values are provided by the power source to the data acquisition system; Q̇ plms is the heat dissipated in the inlet and outlet plenums. Their values were estimated based on the plenums superficial area contacting the fluid, the heat sink average temperature and the local temperature of the refrigerant, using the heat transfer correlations of Stephan and Preußer [62] and Li and Wu [63] for single-phase flow and flow boiling conditions, respectively; Q̇ env is the heat transferred to the environment, calculated as previously described by Chávez et al. [17]. The FFTs were obtained through signal processing in the Origin Lab software and the amplitude of the signals were normalized according to the following equation:

sn =



(3)

The instability parameter was calculated as proposed by Kandlikar et al. [27] and, later, modified by Lee et al. [11]. In this case, the terms related to the channel inlet restriction and associated to a microchannels divergent cross section were neglected. This parameter is given as follows: 5

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Table 1 Uncertainty of measured and calculated parameters. Uncertainty of Parameters Measured

Calculated

Fluid

Hch [µm]

Wch [µm]

Lch [µm]

T °C

p [kPa]

ΔpM [kPa]

ṁ [g/s]

T̅ wall [°C]

T̅ fluid [°C]

G [%]

R600a R290 R1270

22.9

15.4

5

0.15

2 8 8

0.22

0.024

0.4

0.2

13

″ [%] qfp 3.7

and thermocouple in the outlet plenum correspond to 20 and 5 ms, respectively. Thus, these transducers are suitable for the analysis of the phenomena oscillation with frequencies up to 200 Hz. Fig. 5a and b display for R600a the Fast Fourier Transform (FFT) in the frequency domain of the temperature and pressure signals. In these figures, the signal from each transducer for single-phase and diabatic condition is compared with the signal obtained under the presence of convective boiling keeping similar mass velocity, saturation temperature and liquid subcooling at the test section inlet. According to Fig. 5a, the frequencies and the amplitude of the oscillations under conditions of single-phase flow and convective boiling are significantly different. The signal for single-phase flow looks like a white noise with lower amplitude (within the range of uncertainty of temperature measurements) and without a main frequency. In this aspect, Fig. 6 compares images of single flow and flow boiling at the outlet region of the microchannels. These photographic records and the Fig. 5a and b allow to verify the increment of fluctuations with much higher amplitude for the case of the thermocouple located in the output plenum of the device when the phase of the fluid flow changes. This analysis corroborates the existence of a component with an effect on the oscillations which depends on the boiling process and specifically on the growth of bubbles under conditions of confinement. According to Fig. 5b, the pressure oscillation amplitude presents low frequencies of 0.4 Hz for singlephase flow, however, it has a similar amplitude of the pressure oscillations observed under convective boiling conditions with frequencies of 1.4 Hz. In the case of single-phase flow, the frequency peaks may be due to the presence of fluctuations in the liquid flow at the entrance of the microchannels. However, for convective boiling, the frequency peaks may be due to the nucleation intensity of bubbles and the reverse flow. Based on such results, it can be concluded that pressure and temperature oscillations inherent to the phenomena related to flow boiling along the test section are reasonably well segregated from fluctuations associated to the remaining test circuit.

Fig. 5. Normalized temperature and absolute pressure signals for R600a. (a) FFT of the output thermocouple signal, (b) FFT of the absolute pressure signal.

3.1. Effects of heat flux

repeatability of experimental results under single-phase and flow boiling conditions was assured. The K-type thermocouples were calibrated and their uncertainties evaluated according to the procedure proposed by Abernethy and Thompson [66]. The uncertainty of temperature measurements was found equal to 0.15 °C. For the remaining sensors and measuring devices, the uncertainties were assumed equal to the specifications provided by the manufacturers. Accounting for all instrument errors, uncertainties for the calculated parameter were estimated according to the method proposed by Taylor and Kuyatt [67]. Such estimations were performed using EES [68]. The experimental uncertainties associated with the sensors and calculated parameters are given in Table 1.

Figs. 7 and 8 illustrate the effect of the heat flux on the FFT of the temperature signal at the output plenum for R290. According to these figures, the maximum frequency and amplitude of the temperature oscillations increase with increasing the heat flux. This behavior is associated to the intensification of the boiling process due to a higher bubble detachment frequency as observed also in the Fig. 6b. The increment of the temperature amplitude with increasing the heat flux is in agreement with the theory proposed by Kandlikar [27] (see Eqs. (4)–(6)). As shown in Table 2, the force associated to the bubbles growth process, Fback increases as a result of the intensification of the boiling process with the increment of heat flux. This behavior is related to the increase in the mass flow of steam that returns to the input plenum that causes the increment of the oscillation amplitude. On the other hand, values lower than one were calculated for the instability parameter, indicating that no reverse flow of bubbles was noticed for these experimental conditions.

3. Data analysis and discussion This analysis is based on the signals provided by the absolute pressure transducer connected to the inlet of heat sink and the thermocouple located at the output of the microchannels. According to the manufacturers, the response times of the absolute pressure transducer 6

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Fig. 6. Flow at the outlet of the microchannels for R600a, ΔTsub = 5 °C, Tsat = 25 °C, and G = 165 kg/m2 s. (a) Single phase flow; (b) flow boiling q″fp = 99.5 kW/m2. Table 2 Effect of heat flux on forces and instability parameter R. G = 494 kg/m2 s and ΔTsub = 5 °C. q″fp [N/m2]

Fforward [N/m2]

Fback [N/m2]

R [–]

94 165 237 310

495.3 494.0 494.9 495.4

0.063 13.94 51.35 115.5

0.01 0.21 0.32 0.48

Fig. 7. Effect of the heat flux (footprint) on FFT of outlet fluid temperature signal R290, G = 494 kg/m2 s and ΔTsub = 5 °C.

Fig. 9. Effect of the mass velocity on FFT of outlet fluid temperature signal ″ = 280 kW/m2 and ΔTsub = 5 °C. R1270, qfp

inlet plenum for R1270, respectively. In these figures, significant differences between amplitude of the signals are not observed with the variation of mass velocity. However, the oscillation frequency increases considerably as the mass velocity increases. This result is because the increment of the mass velocity favors the rapid detachment of bubbles for a fixed heat flux. Table 3 presents the results obtained for the forces associated to the bubble growth and for the instability parameter with the variation of the mass velocity. In this table a reduction of the back force and of the instability parameter as a result of the increase of the mass velocity is observed. This behavior is explained by the increment of the mass velocity that results in a reduction of the size of the overheated region intensifying the suppression of the nucleation process. Furthermore, if the nucleation sites are active, the increment of the

Fig. 8. Effect of the heat flux (footprint) on FFT of inlet pressure signal R290, G = 494 kg/m2 s and ΔTsub = 5 °C.

3.2. Effects of mass velocity Figs. 9 and 10 illustrate the effect of the mass velocity on the FFT of the temperature in the output plenum and the absolute pressure in the 7

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Fig. 10. Effect of the mass velocity on FFT of inlet absolute pressure signal ″ = 280 kW/m2 and ΔTsub = 5 °C. R1270, qfp

Fig. 12. Effect of liquid subcooling at the inlet test section on FFT of inlet ab″ = 320 kW/m2 and Tsat = 25 °C. solute pressure signal R290, qfp

Table 3 ″ = 280 Effect of mass velocity on forces and instability parameter R. R1270, qfp

Table 4 Effect of liquid subcooling at the inlet test section on forces and instability ″ = 320 kW/m2 and Tsat = 25 °C. parameter R. R290, qfp

kW/m2 and ΔTsub = 5 °C. G [k/m2s]

Fforward [N/m2]

Fback [N/m2]

R [-]

ΔTsub [°C]

Fforward [N/m2]

Fback [N/m2]

R [–]

165 330 494 658

53.83 213.8 482.5 859.2

90.07 74.97 43.07 15.19

1.29 0.59 0.30 0.13

5 10 15

495.6 496.3 495.6

78.08 19.67 0.68

0.40 0.20 0.04

Fig. 13. Effect of saturation temperature on FFT of outlet fluid temperature ″ = 273 kW/m2 and ΔTsub = 10 °C. signal, R290; G = 494 kg/m2 s; qfp

Fig. 11. Effect of liquid subcooling at the inlet test section on FFT of outlet fluid ″ = 320 kW/m2 and Tsat = 25 °C. temperature signal R290, qfp

respectively for R290. In these figures, the oscillation frequency decreases as the liquid subcooling degree at the inlet of the heat sink increases. This result is caused by the sub-cooled conditions of the fluid that reduce the channels length under convective boiling conditions and the average temperature of refrigerantdecreases which favor the suppression of bubble nucleation effects resulting in the reduction of the frequency of detachment thereof. In this aspect, it is highlighted that the nucleation intensity remains constant for a fixed vapor quality. However, with the increment in the degree of subcooling of the liquid at the inlet of microchannels, the temperature gradients in the region where the bubbles collide with the cold fluid are increased originating the increment of the amplitude of the temperature signal. This behavior can be verified in Table 4 in which the values of the boiling forces and instability parameter reduce while the inertial forces remain constant with the increment of the subcooling degree at the inlet heat sink. According to Fig. 12, the variation of liquid subcooling degree has not a significant effect on the amplitude of the pressure oscillations.

mass velocity favors the bubbles detachment with reduced dimensions and decreases the probability of reverse flow. Table 3 presents a value greater than 1.0 for the instability parameter at reduced mass velocities. Already, for mass velocities above 165 kg/m2 s, the instability parameter has values lower than 1.0 which belongs to the stable flow region. This indicates that reverse flow of bubbles could exist for reduced mass velocities and consequently increase the effects of thermal instabilities. This behavior was also verified by Lee et al. [11] and Leão et al. [16] and stems from the superposition of boiling forces over the inertial ones. Similar overall trends for the influence of mass velocity on temperature oscillations were found for the other two refrigerants. 3.3. Effects of liquid subcooling at the inlet of test section Figs. 11 and 12 illustrate the effect of the degree of liquid subcooling at the inlet plenum on the FFT of the temperature signal at the output plenum and absolute pressure transducer at the inlet plenum, 8

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Fig. 14. Effect of saturation temperature on FFT of outlet fluid temperature ″ = 250 kW/m2 and ΔTsub = 10 °C. signal, R1270; G = 494 kg/m2 s; qfp

Fig. 16. Effect of fluid on FFT of outlet fluid temperature signal, G = 494 kg/ ″ = 280 kW/m2 and ΔTsub = 5 °C. m2 sqfp

Fig. 15. Effect of saturation temperature on FFT of inlet absolute pressure ″ = 273 kW/m2 and ΔTsub = 10 °C. signal, R290; G = 494 kg/m2 s; qfp

Fig. 17. Effect of fluid on FFT of inlet absolute pressure signal, G = 494 kg/ ″ = 280 kW/m2 and ΔTsub = 5 °C. m2 sqfp

Table 5 Effect of fluid on forces and instability parameter R. G = 494 kg/m2 s; ″ , R1270 = 250 kW/m2 and ΔTsub = 10 °C. ″ , R290 = 273 kW/m2; qfp qfp

Table 6 ″ = 280 Effect of fluid on forces and instability parameter R. G = 494 kg/m2 sqfp kW/m2 and ΔTsub = 5 °C.

Fluid

Tsat [°C]

Fforward [N/m2]

Fback [N/m2]

R [–]

Fluid

Fforward [N/m2]

Fback [N/m2]

R [–]

R290 R290 R1270 R1270

21 25 21 25

486.4 496.3 476.6 482.7

22.83 19.67 13.75 12.48

0.22 0.20 0.17 0.16

R600a R290 R1270

442.9 494.9 482.5

131.4 51.35 43.07

0.54 0.32 0.30

21 °C to 25 °C produces an increase of the forward force and a decreased back force which result in a slight variation of the instability parameter. Such behavior may be associated to the increment of the vapor density generated by the increased value of the saturation temperature as can be analyzed in the Eqs. (4)–(6). In this aspect, the increment of nucleation intensity combined with the increase of inertial forces cause the elevation of the amplitude oscillations when the saturation temperature is increasing.

3.4. Effects of saturation temperature Figs. 13–15 describe the effect of the saturation temperature on the FFT of the temperature signal at the output plenum and absolute pressure signal at the inlet plenum, respectively. According to these figures, the higher oscillation frequencies are observed for the lower saturation temperatures. This variation is related to the increment of the latent heat of vaporization when the saturation temperature of the refrigerant decreases. In this condition, the fluid present higher vapor phase specific volume which result in an increment of the frequency for bubbles detachment. On the other hand, an increase in the amplitude of the temperature and pressure oscillations at higher saturation temperatures is observed in the Figs. 13–15. This behavior is due to the increment of nucleation intensity when the saturation temperature increases. Table 5 shows the results of the effect of saturation temperature on forces associated with bubble growth and the instability parameter. According to this table, the elevation of saturation temperature from

3.5. Effects of refrigerant fluid Figs. 16 and 17 illustrate the effect of the refrigerant fluid on the FFT of the temperature signal at the outlet plenum and inlet absolute pressure signal, respectively. According to these figures, the oscillation frequencies present the following order of increment R1270, R290 and R600a. This behavior is due to the higher value of the specific volume of the R600a vapor that increase the vapor velocities with an increment of the frequency oscillations for both temperature and pressure. 9

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Furthermore, an analysis based on the Clapeyron equation allows to verify the effect of the fluid properties of R600a, R290 and R1270 on the frequency of temperature and pressure oscillations. A simple calculation of the gradient (dp/dT)sat for the three refrigerant operating under the same condition gives lower values for R600a, intermediate values for R290 and a higher gradient for R1270, being in agreement with the measured frequency of the oscillations. This analysis allows to conclude that the fluid with a higher ratio between the latent heat of vaporization and the product between the absolute temperature and the specific volume of evaporation has better operating conditions with lower oscillation frequencies. Fig. 16 shows that the operation with R290 has higher temperature oscillation amplitudes and similar magnitudes for the other fluids. Such behavior can be explained because R290 has a higher latent heat of vaporization and lower surface tension. These characteristics favors the velocity of bubble detachment and can cause the increase in the amplitude of the thermal oscillations. Table 6 presents the results for the fluid refrigerant effect on the forces associated to bubble growth and to the values of the instability parameter. According to this table, R290 shows higher values of liquid inertial forces. This issue can be related to the larger amplitude of the oscillations presented by R290 illustrated in Figs. 16 and 17. In this case, the liquid R290 enters the microchannels with greater force at a lower pressure and temperature and then impacts with the reverse vapor flow and consequently increases the amplitude of the oscillations. On the other hand, R1270 has lower values for the back force and the R600a presents higher values of the instability parameter which could result in a higher probability of reverse flow. Note that the instability parameter has the same incremental order as the oscillation frequencies shown in Fig. 16 (R1270/R290/R600a). According to Figs. 16 and 17, R290 has the higher oscillation peak of temperature and pressure, while R1270 and R600a have oscillation peaks with similar magnitudes but at different frequencies at 2.4 and 3.8 Hz, respectively. In addition, the R1270 presents other oscillation peaks at frequencies of 2 Hz and 6.2 Hz, in this case, slightly higher than the amplitude observed for the other fluids. However, based on the results displayed in these figures and in contrast to the values of the parameter of instability reported in Table 6, it is possible to conclude that the R1270 fluid presents lower thermal oscillations because its signal does not have the highest peak and presents the lowest instability parameter compared to the other two refrigerants.



Acknowledgements The authors acknowledge Comisión Nacional Científica y Tecnológica, CONICYT-Chile for support received in the Fondo Nacional de Desarrollo Científico y Tecnológico, FONDECYT, project N°3170214. References [1] F.T. Kanizawa, C.B. Tibiriçá, G. Ribatski, Heat transfer during convective boiling inside microchannels, Int. J. Heat Mass Transf. 93 (2016) 566–583. [2] F. Ronshin, E. Chinnov, Experimental characterization of two-phase flow patterns in a slit microchannel, Exp. Therm. Fluid Sci. 103 (2019) 262–273. [3] S.G. Kandlikar, A general correlation for saturated two-phase flow boiling heat transfer inside horizontal and vertical tubes, J. Heat Transf. 112 (1990) 219–228. [4] G. Hetsroni, A. Mosyak, Z. Segal, Nonuniform temperature distribution in electronic devices cooled by flow in parallel microchannels, Comp. Packag. Technol., IEEE 24 (2001) 16–23. [5] W. Qu, I. Mudawar, Prediction and measurement of incipient boiling heat flux in micro-channel heat sinks, Int. J. Heat Mass Transf. 45 (2002) 3933–3945. [6] B. Daniels, J.A. Liburdy, D.V. Pence, Adiabatic flow boiling in fractal-like microchannels, Heat Transf. Eng. 28 (2007) 817–825. [7] R.R. Muwanga, R.R. MacDonald, I. Hassan, Characteristics of flow boiling oscillations in silicon microchannel heat sinks, J. Heat Transf. 129 (2007) 1341–1351. [8] D. Bogojevic, K. Sefiane, A.J. Walton, G.C.H. Lin, Two-phase flow instabilities in microchannels, ECI International Conference on Heat Transfer and Fluid Flow in Microscale, Whistler, Canada, (2008). [9] T. Harirchian, S.V. Garimella, Microchannel size effects on local flow boiling heat transfer to a dielectric fluid, Int. J. Heat Mass Transf. 51 (2008) 3724–3735. [10] B. Agostini, J.R. Thome, M. Fabbri, B. Michel, D. Calmi, U. Kloter, High heat flux flow boiling in silicon multi-microchannels - part I: heat transfer characteristics of refrigerant R236fa, Int. J. Heat Mass Transf. 51 (2008) 5400–5414. [11] H.J. Lee, D.Y. Liu, S. chune Yao, Flow instability of evaporative micro-channels, Int. J. Heat Mass Transf. 53 (2010) 1740–1749. [12] E. Costa-Patry, J. Olivier, B.A. Nichita, B. Michel, J.R. Thome, Two-phase flow of refrigerants in 85 μm-wide multi-microchannels: Part I - pressure drop, Int. J. Heat Fluid Flow 32 (2011) 451–463. [13] F.J.L. Do Nascimento, Hugo Leonardo Souza Lara, Ribatski, Gherhardt, An experimental study on flow boiling heat transfer of r134a in a microchannel-based heat sink, Exp. Therm. Fluid Sci. 45 (2013) 117–127. [14] S. Szczukiewicz, N. Borhani, J.R. Thome, Fine-resolution two-phase flow heat transfer coefficient measurements of refrigerants in multi-microchannel evaporators, Int. J. Heat Mass Transf. 67 (2013) 913–929. [15] H. Lee, I. Park, I. Mudawar, M.M. Hasan, Micro-channel evaporator for space applications – 1. Experimental pressure drop and heat transfer results for different orientations in earth gravity, Int. J. Heat Mass Transf. 77 (2014) 1213–1230. [16] H.L.S.L. Leão, C.A. Chávez, F.J. do Nascimento, G. Ribatski, An analysis of the effect of the footprint orientation on the thermal-hydraulic performance of a microchannels heat sink during flow boiling of r245fa, Appl. Therm. Eng. 90 (2015) 907–926. [17] C.A. Chávez, H.L.S.L. Leão, G. Ribatski, Evaluation of thermal-hydraulic performance of hydrocarbon refrigerants during flow boiling in a microchannels array heat sink, Appl. Therm. Eng. 111 (2017) 703–717. [18] D. Bogojevic, K. Sefiane, A.J. Walton, H. Lin, G. Cummins, D.B.R. Kenning, T.G. Karayiannis, Experimental investigation of non-uniform heating effect on flow boiling instabilities in a microchannel-based heat sink, Int. J. Therm. Sci. 50 (2011) 309–324. [19] G. Ribatski, A critical overview on the recent literature concerning flow boiling and two-phase flows inside micro-scale channels, Exp. Heat Transf. 26 (2013) 198–246. [20] C.B. Tibiriçá, L.E. Czelusniak, G. Ribatski, Critical heat flux in a 0.38 mm microchannel and actions for suppression of flow boiling instabilities, Exp. Therm. Fluid Sci. 67 (2015) 48–56. [21] J.A. Boure, A.E. Bergles, L.S. Tong, Review of two-phase flow instability, Nucl. Eng. Des. 25 (1973) 165–192. [22] S. Kakaç, T.N. Veziroglu, M.M. Padki, L.Q. Fu, X.J. Chen, Investigation of thermal instabilities in a forced convection upward boiling system, Exp. Therm. Fluid Sci. 3 (1990) 191–201. [23] W. Qu, I. Mudawar, Measurement and correlation of critical heat flux in two-phase micro-channel heat sinks, Int. J. Heat Mass Transf. 47 (2004) 2045–2059. [24] W. Qu, I. Mudawar, Flow boiling heat transfer in two-phase micro-channel heat sinks. Part I. Experimental investigation and assessment of correlation methods, Int. J. Heat Mass Transf. 46 (2003) 2755–2771. [25] J. Lee, I. Mudawar, Two-phase flow in high-heat-flux micro-channel heat sink for refrigeration cooling applications: Part I – pressure drop characteristics, Int. J. Heat

4. Conclusions In the present work a new database was obtained concerning experimental data for thermal oscillations of R290, R600a and R1270 refrigerants during flow boiling in a microchannel array heat sink. The database was parametrically analyzed and based on the present study; the following main conclusions are drawn:

• The FFT analysis allows to evaluate the effect of thermal instabilities





velocity does not have a significant effect on either amplitude of temperature or pressure oscillations. The increment of the frequency of oscillations was found for increasing heat flux and mass velocity, decreasing degree of liquid subcooling at the inlet of the test section and decreasing saturation temperature.

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