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Critical heat flux characteristics of R1234yf, R1234ze(E) and R134a during saturatedflow boiling in narrow high aspect ratio microchannels

Kærn, Martin Ryhl; Criscuolo, Gennaro; Meyer, Knud Erik; Markussen, Wiebke Brix

Published in:International Journal of Heat and Mass Transfer

Link to article, DOI:10.1016/j.ijheatmasstransfer.2021.121840

Publication date:2021

Document VersionPublisher's PDF, also known as Version of record

Link back to DTU Orbit

Citation (APA):Kærn, M. R., Criscuolo, G., Meyer, K. E., & Markussen, W. B. (2021). Critical heat flux characteristics ofR1234yf, R1234ze(E) and R134a during saturated flow boiling in narrow high aspect ratio microchannels.International Journal of Heat and Mass Transfer, 180, [121840].https://doi.org/10.1016/j.ijheatmasstransfer.2021.121840

International Journal of Heat and Mass Transfer 180 (2021) 121840

Contents lists available at ScienceDirect

International Journal of Heat and Mass Transfer

journal homepage: www.elsevier.com/locate/hmt

Critical heat flux characteristics of R1234yf, R1234ze(E) and R134a

during saturated flow boiling in narrow high aspect ratio

microchannels

Martin Ryhl Kærn

∗, Gennaro Criscuolo , Knud Erik Meyer , Wiebke Brix Markussen

Department of Mechanical Engineering , Technical University of Denmark , Nils Koppels Allé, Building 403, 2800 Kongens Lyngby , Denmark

a r t i c l e i n f o

Article history:

Received 10 March 2021

Revised 18 June 2021

Accepted 9 August 2021

Keyword:

Critical heat flux

Microchannels

Flow boiling

Low-GWP refrigerant

Electronics cooling

a b s t r a c t

This experimental study investigated narrow and high aspect ratio multi-microchannels that were micro-

machined in copper with thin separating walls during saturated flow boiling of refrigerants. The hypothe-

sis was that these channels could increase the footprint critical heat flux and support the future develop-

ment of thermal management systems for power electronics and other electronic packages. The measured

footprint critical heat flux was as high as 678.5 W/cm

2 , which is twice as high as other investigations in

the literature concerning saturated flow boiling of refrigerants. Two test samples were fabricated with a

footprint area of (10 × 10) mm. The first had 25 rectangular channels (198 μm wide, 1167 μm high).

The second had 17 rectangular channels (293 μm wide, 1176 μm high). The experimental investigation

covered low-GWP replacement refrigerants (R1234yf, R1234ze(E)) as well as the well-examined R134a

serving as a benchmark. A large data bank was obtained with 432 data points covering a wide range of

typical inlet subcooling (1.3–14.7) K and mass fluxes (333–1260) kg/m

2 s as well as two nominal satura-

tion temperatures (30 and 40) °C.

The effect of inlet subcooling was found to be consistently significant at the higher mass fluxes. This

was contradictory to other investigations that found moderate or insignificant effects. The result is sug-

gested to be attributed to the two-phase stability induced by the upstream throttle valve, inlet orifices,

isolated refrigerant in the inlet and outlet plenums as well as the short heated length causing higher im-

portance of orifice to channel pressure drop importance. Finally, a new modified Katto and Ohno correla-

tion including the effect of subcooling was proposed, achieving a 4.0% mean average error and predicting

93.3% of the data points within 10% error.

© 2021 The Authors. Published by Elsevier Ltd.

This is an open access article under the CC BY license ( http://creativecommons.org/licenses/by/4.0/ )

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. Introduction

The continuous miniaturization and advancement in the perfor-

ance of modern microelectronics have resulted in a sustained in-

rease of the cooling demands and the maximum heat flux targets

f applications such as high-performance computing and power

lectronics. The higher heat flux targets must be provided while

aintaining the material temperatures below prescribed limits.

he current target for power electronics in electric vehicles is

lose to 250 W/cm

2 at the chip level, while in some niche ap-

lications, such as defense power electronics, it can reach up to

Abbreviations: CHF, critical heat flux; GWP, global warming potential; MAE,

ean average error; ODP, ozone depletion potential; RTD, resistance thermal de-

ector. ∗ Corresponding author.

E-mail address: pmak@mek.dtu.dk (M.R. Kærn).

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017-9310/© 2021 The Authors. Published by Elsevier Ltd. This is an open access article u

0 0 0 W/cm

2 [1] . Flow boiling in microchannels represents one of

he most promising cooling solutions for these applications, due to

1) the high heat transfer during boiling, (2) the large surface area

o volume ratio of microchannel geometries and (3) low pump-

ng power [ 2 , 3 ]. The maximum achievable heat flux is generally

eferred to as the critical heat flux (CHF). Surpassing the CHF com-

only results in a sharp increase of the heated surface tempera-

ure and may lead to a physical burnout of the electronics. Two

hysical mechanisms are commonly associated with the CHF, de-

ending on whether subcooled flow boiling ( x out < 0 ) or saturated

ow boiling ( x out > 0 ) occurs. For subcooled flow boiling, the liq-

id replenishment of the heat transfer surface is hindered by the

apor formation during nucleate boiling (departure from nucleate

oiling, DNB). For saturated flow boiling, the large increase in void

raction trigger annular flow in which dryout of the liquid film is

he main CHF mechanism [ 1 , 2 ]. Two-phase flow instabilities in mi-

rochannels may trigger earlier CHF such as upstream compress-

nder the CC BY license ( http://creativecommons.org/licenses/by/4.0/ )

M.R

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Fluids Material Conditions Effects

L h /D he H ch /W ch

T sat ,

G �T sub ,

q ′′ f p,max

G �T sub q fp,max

] [-] [-] [ °C] [kg/m

2 s] [K] [W/cm

2 ]

131.2 0.15 R-123 Copper 42–55 411–534 24–38 16.4 ∗ CHF ↑ with G

344.7 1.32 R-123 Silicon 52–82 291–1118 29–59 ∗∗∗∗ 196 CHF ↑ with G, �T sub CHF ↑↓ with T sat

52.2 3.0 R236fa Silicon 20–34 276–992 0–15 250 CHF ↑ with G CHF → with �T sub , T sat

22.6 8.7 R245fa, R236fa,

R134a

Copper 10–50 90–450 0–24 342 CHF ↑ with G, �T sub CHF → with T sat

56.9 3.8 15–40 200–4000 0–15 215 CHF ↑ with G CHF → with �T sub , T sat

28.4 3.8 R245fa, R236fa,

R134a

Copper 15–40 250–1500 5–25 330 CHF ↑ with G, �T sub ∗∗ CHF ↓ with T sat

∗∗

71.7 3.9 FC-77 Silicon 97 254–1015 26 107 CHF ↑ with G

0 (12.5–

17.5)

0.5 R134a,

R1234ze(E),

R1234yf, R32

Aluminum 25–76 146–1504 0–18 175 ∗ CHF ↑ with G CHF → with T sat ∗∗∗

CHF ↓↑ with L h / D he ∗∗∗

3 26.3 0.25 25–76 146–1504 0–18 169 ∗

3 87.5 0.2 Ethanol,

acetone

Copper 56–78 3–39 29–44 68 CHF ↑ with G CHF ↓ with D he

116.7 0.2 56–78 5–70 29–44 90

73.6 1.23 R134a Copper 20–28 x in 1000 x in = 0.01–

0.20

71 ∗ CHF ↑ with G CHF ↓ with x in , T sat

0 3.9 - R113 Copper 57 31–150 10–32 200 CHF ↑ with G, D CHF → with �T sub

19.6 - 57 120–480 10–32 256

increase

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Table 1

Main parameters, conditions and effects observed in multi-microchannel satura

Authors Geometry

N × ( W ch × H ch ) ,

L h D he , L h / D he , H ch / W ch L h D he

[ μm] [mm] [ μm

Kuan and Kandlikar [28] (2006) 6 × (1054 × 157) 63.5 484

Ko ̧s ar and Peles [13] (2007) 5 × (200 × 264) 100 290

Agostini et al. [14] (2008) 67 × (223 × 680) 20 383

Park and Thome [6] (2010) 20 × (467 × 4052) 20 883

29 × (199 × 756) 20 352

Mauro et al. [12] (2010) 29 × (199 × 756) 10 352

Chen and Garimella [29] (2012) 60 × (100 × 389) 12.7 177

Mastrullo et al. [16] (2017) 7 × (2000 × 1000) 25–35 200

7 × (2000 × 500) 35 133

Hong et al. [17] (2018) 4 × (2000 × 400) 100 114

4 × (1500 × 300) 100 857

Dalkılıç et al. [15] (2020) 27 × (382 × 470) 40 543

Bowers and Mudawar [7] (1994) 3 (circular channels) 10 254

17 (circular channels) 10 510

∗ maximum footprint CHF ( q ′′ f p,max

) calculated by assuming 0.9 fin efficiency ↑ ∗∗ no effect for R245fa → no effect ∗∗∗ CHF decreased with T sat for R134a ↓ decrease ∗∗∗∗ based on 23 °C room/inlet temperature

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

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hannel CHF investigations have been omitted as well as the re-

earch works using water as the working fluid.

The studies summarized in Table 1 cover both shallow

H ch / W ch < 1 ) and narrow ( H ch / W ch > 1 ) rectangular multi-

hannels as well as a single circular multichannel study. The

eated equivalent diameters (based on 3-sided heating for the

ectangular channels) range from 177 μm to 2.54 mm with most

alues below 1 mm. The heated length ranges from 10 mm to

00 mm and results in length-to-diameter ratios from 4 to 350,

ith most values concentrated in the range from 10 to 60. The

ulti-microchannels are fabricated in silicon, copper and alu-

inum, and a wide range of fluids have been tested, including a

ingle Fluorinert electronic coolant, ethanol, acetone and several

efrigerants. Most of the refrigerants have been phased out or

re planned to be phased out in the coming years due to their

igh global warming potential (GWP) or Ozone Depletion Potential

ODP). The operational conditions cover a large range of saturation

emperatures, mass fluxes and inlet subcooling. The maximum

HF based on the footprint area is also reported and has been

alculated by assuming a fin efficiency of 0.9 in the studies marked

ith asterisk “∗”. These studies reported channel wall CHF solely.

s can be seen from Table 1 , the maximum footprint CHF obtained

n previous studies are only partly covering current targets for

ower electronics in electric vehicles. In this context, it should be

oticed that prior studies did not always attempt to achieve a high

HF, as pointed out by Park and Thome [6] .

.1. Effect of parameters

Several geometric parameters and fluid conditions affect the

HF in multi-microchannel flow boiling, i.e. channel size, heated

ength, surface roughness, fluid, mass flux, inlet subcooling, satura-

ion temperature. Some of the main conclusions regarding the ef-

ect of these parameters are also included in Table 1 . Generally, the

bservations indicate that the CHF increases with increasing mass

ux.

Furthermore, the CHF either increases or is unaffected by the

evel of inlet subcooling. Bowers and Mudawar [7] attributed the

ndependence of inlet subcooling to the relatively short distance of

he subcooled region in their heated section. Later, Qu and Mu-

awar [8] observed similar independence in their multi-channel

xperiments with water. They attributed the independence of in-

et subcooling to instability inherent to the multi-channel con-

guration, where periodic vapor backflow into the inlet header

ixed with the incoming subcooled liquid. However, even in sin-

le microchannel CHF investigations adopting an upstream stabi-

izing valve, the inlet subcooling was found to be independent [9–

1] . Park and Thome [6] found the effect of subcooling to be mod-

rate for their larger diameter multi-channels, while their smaller

iameter channels did not indicate such an effect. Using the same

maller diameter channels and adopting a liquid supply split sys-

em that reduced the heated length to half, Mauro et al. [12] found

hat an increasing subcooling did have an increasing effect on the

HF for refrigerant R134a and R236fa, but no effect was observed

or R245fa. The above findings could suggest that the effects of

ubcooling on the CHF are highly linked to the flow instabilities in

ulti-channels, the channel size, the heated length and the fluid.

o ̧s ar and Peles [13] observed an increasing effect of subcooling on

HF, however, these experiments were also carried out at very high

ubcooling. On the other hand, Agostini et al. [14] did not observe

ignificant effects on the level of subcooling in their small diam-

ter channels. Concerning the experiments carried out by Dalkılıç

t al. [15] with saturated inlet conditions, the CHF was found to

ecrease with increasing inlet vapor quality.

Similar inconsistencies are found in the literature concerning

he effects of saturation temperature on the CHF as well as its

3

ignificance. Ko ̧s ar and Peles [13] found both increasing and de-

reasing trends concerning the saturation temperature, indicating

peak CHF. Park and Thome [6] found increasing and decreasing

rends for their larger diameter channels and their smaller diam-

ter channels, respectively. These effects were suggested to be in-

ignificant. Mauro et al. [12] and Dalkılıç et al. [15] found a de-

rease in CHF with increasing saturation temperature, while Agos-

ini et al. [14] and Mastrullo et al. [16] found an insignificant effect

r a slightly increasing and decreasing trend. The increasing and

he decreasing trend is commonly explained by the counteracting

ffects related to the refrigerant properties with increasing satura-

ion temperature. On one hand, the increased vapor to liquid den-

ity ratio tends to stabilize the liquid film during annular flow and

educes the liquid entrainment into the vapor core, thus increas-

ng the CHF. On the other hand, the reductions in the latent heat

f vaporization and the surface tension lead to a reduced capacity

nd a more breakable liquid film [16] , respectively, thus decreasing

he CHF.

Concerning the geometric parameters such as the heated equiv-

lent diameter and the heated length, it is difficult to make a di-

ect comparison regarding the CHF in multi-channels since most

tudies involve only a single test sample. The few studies with

ore diameters, i.e. Bowers and Mudawar [7] and Park and Thome

6] , indicate that the channel wall CHF increases with increas-

ng diameter at the same mass flux. This result is consistent with

ingle-channel CHF investigations. On the other hand, Hong et al.

17] found a decrease in channel wall CHF with increasing diam-

ter in their shallow microchannel test samples. The effect of the

eated length can be examined by comparing the small diameter

esults of Park and Thome [6] with the liquid supply split sys-

em adopted by Mauro et al. [12] for the same channels. Mauro

t al. [12] report a CHF increase of up to 80% on the footprint

evel. This is also consistent with single-channel CHF investigations

e.g. Wojtan et al. [9] ), which indicate an increase in the CHF with

ecreasing heated length. Concerning the length-to-diameter ratio,

he CHF typically decreases with increasing length-to-diameter ra-

io, which was also found at low mass fluxes in the investigation

y Mastrullo et al. [16] . However, at higher mass fluxes their re-

ults indicated the opposite trend.

.2. Purpose, hypothesis and objectives

The purpose of the current investigation was to achieve a

igher footprint CHF by employing narrow high aspect ratio multi-

icrochannels during saturated flow boiling of refrigerants. The

ypothesis is that narrow high aspect ratio channels with thin sep-

rating walls may increase the heat transfer performance in terms

f footprint CHF, and support the future development of ther-

al management systems for power electronics and other elec-

ronic packages. In this context, low-GWP refrigerants (R1234yf,

1234ze(E)) and the traditional refrigerant R134a, serving as a

enchmark, were considered viable working fluids for the purpose

f the study. To test the hypothesis, two multi-microchannels test

amples having a footprint area of (10 × 10) mm were fabricated

n copper by micro-milling. The first test sample had 25 channels,

ach 200 μm wide, 1200 μm high and 200 μm separating walls,

ominally. The second test sample had 17 channels, each 300 μm

ide, 1200 μm high with 300 μm separating walls, nominally.

hese values translate into aspect ratios of 4 and 6, heated equiv-

lent diameters of 369 μm and 533 μm, and length-to-diameter

atios of 19 and 27, respectively. Effective channel dimensions were

easured in a confocal microscope and presented in Table 2 .

The novelty of the study consisted in obtaining CHF data in nar-

ow and high aspect ratio multi-microchannels with low-GWP re-

rigerants. Such an investigation was missing according to the lit-

rature survey presented in Table 1 . Furthermore, the study aimed

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Table 2

Channel dimensions and standard deviations measured by a confocal microscope.

Label W ch H ch W f in D he L h / D he H ch / W ch N

[ μm] [ μm] [ μm] [ μm] [-] [-] [-]

Geometry “F” 198 ± 6.73 1167 ± 1.43 200 ± 8.08 365 27.4 5.89 25

Geometry “D” 293 ± 6.47 1176 ± 0.98 306 ± 8.20 521 19.2 4.01 17

Fig. 1. Schematic diagram of the experimental flow boiling test facility.

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o provide even higher footprint CHF levels, thus scientifically con-

olidating the suitability of refrigerants towards the future elec-

ronics cooling need. The objectives of the study were formulated

s follows: the first objective was to document an improvement

f footprint CHF by using narrow and high aspect ratio multi-

icrochannels. The second objective was to investigate the effects

f the main parameters on the CHF (mass flux, inlet subcooling,

aturation temperature, heated equivalent diameter). The third ob-

ective was to evaluate the applicability of existing CHF correla-

ions for predicting the CHF in the investigated narrow and high

spect ratio channels with low-GWP refrigerants.

.3. Outline

The manuscript includes a description of the experimental

ethods in Section 2. This description includes the test rig, the

est section and the CHF detection method. Furthermore, the data

eduction method and the experimental uncertainties involved are

eported as well as the experimental design points. The results are

resented in Section 3 in terms of both footprint CHF and channel

all CHF, including an analysis of the effects of the main parame-

ers. The CHF results are compared with existing CHF correlations

nd a new modified Katto and Ohno correlation accounting for the

nlet subcooling is presented. Finally, the results are discussed in

ection 4 and followed up by the conclusions in Section 5.

. Experimental methods

.1. The refrigerant loop

A schematic of the experimental test facility is illustrated in

ig. 1 . A Tuthill gear pump circulated the refrigerant in the main

oop and was operated digitally using a variable speed drive. From

he receiver, the liquid was circulated through a plate heat ex-

hanger (subcooler) and a filter drier, and pumped downstream

hrough a Micro Motion Coriolis Elite sensor (CMFS010M), a pre-

eater, a flow balancing needle valve and a 7-micron filter be-

ore entering the test section. After the test section, a large plate

4

eat exchanger condensed the evaporated refrigerant with mini-

al pressure drop and eventually returned the refrigerant to the

eceiver. The saturation pressure in the system was controlled by

n internal helical coil inside the receiver, in which the water

ow temperature was controlled. This was achieved by using a

hree-way valve for the cooling as well as a heating wire wrapped

round the piping and connected to a thyristor power controller

or the heating. Three-way valves were also used to control the

ater temperature flowing through the condenser and the sub-

ooler. The preheater consisted of a heating wire wrapped around

he refrigerant piping and operated using a Keysight DC power

upply (N5749A). To avoid spurious pulsations generated by the

ump through the piping to the mass flow meter, a flexible hose

as installed between these two components. The test-section by-

ass was closed during all measurements, and so was the high-

ressure solenoid safety valve. Danfoss AKS33 pressure transmit-

ers (060G2044 and 060G2050) and Omega Engineering type-T

hermocouples were placed at several locations in the flow loop

s well as in the test section to monitor the refrigerant thermody-

amic conditions. National Instruments CompactDAQ modules and

abview were used to record all the data as well as to control tem-

eratures, saturation pressure and power supplies. The tempera-

ure sensors were calibrated by using a thermal bath and a Rose-

ount Standard Platinum Resistance (model 162), while the pres-

ure sensors were calibrated by using a Budenberg Air-Operated

eadweight tester. Table 3 reports the instruments’ ranges and the

ncertainties based on the 95% confidence interval.

.2. Test section

The test section is illustrated using a schematic cross-sectional

rawing in Fig. 2 and an exploded view of the CAD model as-

embly in Fig. 3 . The test section was designed to facilitate inter-

hangeable microchannel heat sinks. In the current investigation,

wo heat sinks were employed: geometry “D” was the heat sink

ith the (300 × 1200) μm nominal channels, and geometry “F”

as the heat sink with the (200 × 1200) μm nominal channels.

he microchannel heat sinks were manufactured by micro-milling

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 2. Schematic drawing of the test section (not to scale).

Table 3

Instruments ranges and uncertainties based on 95% confidence.

Instrument Uncertainty Unit Range

Thermocouples,

T-type

± 0.15 °C 18–50

Thermocouples,

K-type

± 0.16 °C 18–50

RTDs ± 0.09 °C 25–160

Low pressure

transmitter

± 0.025 bar 0–10

High pressure

transmitter

± 0.04 bar 0–21

Mass flow meter ± 0.1% kg/hr 5–55

Power supply (test

section)

± 0.1% + 0.15 V 0–150

± 0.1% + 0.03 A 0–10

Power supply

(preheater)

± 0.1% + 0.1 V 0–100

± 0.1% + 0.0225 A 0–7.5

Differential

pressure

transmitter

± 0.19% mbar 0–500

i

f

w

O

t

r

f

r

s

t

t

m

n

s

t

a

a

P

t

p

f

a

o

Fig. 3. Exploded view of the CAD model assembly. (1) Stainless steel cover, (2)

Teflon plate, (3) Cover glass, (4) PEEK inserts with and w/o. orifices, (5) Copper heat

sink with microchannels, (6) Insulating PEEK inserts, (7) Stainless steel chassis, (8)

Serpentine heater.

s

a

u

t

i

(

i

t

t

c

g

b

T

n Electrolytic Tough Pitch Copper (C101/CW004A). After manu-

acturing, the effective dimensions of the channels and separating

alls were measured using a confocal microscope (Olympus Lext

LS40 0 0) with a dedicated 20x lens, reported in Table 2 . Similarly,

he surface roughness was measured using identical processed sur-

ogate channels, which allowed for the opening of the channels

or the surface roughness inspection. For geometry “F”, the average

oughness values were 372 nm and 348 nm for the climb-milled

ide and the conventional-milled side, respectively. Likewise, for

he geometry “D”, the values were 662 nm and 504 nm, respec-

ively. More information about the characterization of the multi-

icrochannels is presented in Criscuolo et al. [18] .

A stainless steel chassis featured the fluid inlet and outlet con-

ections to the microchannel heat sinks, as well as the pres-

ure taps and 2-by-2 thermocouples placed through the bottom of

he steel structure at the inlet and the outlet, respectively. Both

bsolute pressure transmitters (the low-pressure range, Table 3 )

nd differential pressure transmitter (Endress-Hauser Deltabar S

MD75) were connected to the pressure taps, and thermocouples

ype-K with a diameter of 0.25 mm were used. PEEK inserts were

laced at the inlet and the outlet plenum to isolate the fluid flow

rom the copper heat sink in locations without the microchannels

s well as to provide inlet orifices to the microchannels. The inlet

rifices had flow area restrictions of (35 and 40) % of the mea-

5

ured channel flow area for geometry “D” and “F”, respectively,

nd the orifice length was 5 mm. Additional PEEK inserts were

sed to minimize the heat losses from the copper heat sink to

he stainless steel cover and the stainless steel chassis. A borosil-

cate glass window gave visual access with a high-speed camera

Photron Nova S9 with a double 2x teleconverter) and pulsed light-

ng (2 × 150 W GSVitec LED lights). A 100 μm heat resistant and

ransparent silicon layer was placed between the glass window and

he microchannels to avoid boiling on top of the fins as well as

hannel-to-channel leakage flows.

A platinum serpentine heater (10 × 10 mm) with four inte-

rated 4-wire resistance thermal detectors (RTDs) were fabricated

y vapor deposition (883 nm high) on a 350 μm silicon wafer.

he heaters were vacuum soldered onto the bottom of each copper

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 4. CHF detection example (accelerated voltage increments). RTD0 to RTD3 rep-

resent the RTDs placed at 7/8, 5/8, 3/8 and 1/8 of the heated length in the flow

direction, respectively.

h

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f

5

f

f

d

t

p

w

d

d

o

c

h

A

i

t

2

v

R

a

b

b

n

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s

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i

d

a

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t

6

a

6

a

T

t

o

2

c

T

a

s

t

a

p

g

d

b

h

q

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c

m

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t

q

η

m

w

fi

h

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T

R

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w

b

l

eat sink using Sn-Ag solder paste placed in a 20 μm deep groove

ith a similar area as the chip heater. Supports were employed

o ensure the vertical alignment of the serpentine heater with the

icrochannel footprint area during the soldering. The solder lay-

rs were checked for significant voids using 3-D tomography and

ound to be less than 1%. The shrinkage of the solder paste was

5%, the melting point 221 °C and the thermal conductivity data as

unction of temperature was obtained from the manufacturer. Dif-

erential height measurements confirmed the shrinkage of the sol-

er. Another solder type with a melting point at 183 °C was used

o connect the heater pads with connectors for the Keysight DC

ower supply (N5770A). A maximum heater temperature of 165 °C

as considered during the experiments to avoid melting of the sol-

er. When this temperature was exceeded by one of the RTDs, a

igital output signal from Labview triggered a safe-stop function

f the power supply and ensured a safe operation of the test fa-

ility. The serpentine heater was painted black with a black matt

eat-resistant paint to allow for infrared thermography with FLIR

655sc and a 50 μm close-up lens. Finally, the RTDs were cal-

brated with a Rosemount Standard Platinum Resistance using a

hermal oil bath.

.3. CHF detection

The CHF detection was performed by gradually increasing the

oltage of the power supply by 0.5 V until a sharp increase in the

TDs occurred, which immediately triggered the safety function

nd cut-off the power supply. Fig. 4 exemplifies the CHF detection

y an accelerated detection in which the voltage was incremented

y 5 × 5 V, 3 × 1 V and 1 × 0.5 V. When the CHF condition was

ot triggered, the thermal response of the RTDs indicated a first-

rder response. When the critical heat flux occurred, the thermal

esponse of the RTDs indicated a very sharp growth. In almost any

ase, the RTD placed closest to the channel exit reached the high-

st temperature. At the lowest mass flux values, the thermal re-

ponse was not sharply increasing during CHF. In fact, raising the

oltage further in these conditions did not increase the power up-

ake because the resistance of the platinum heater increased with

ncreasing temperature. Therefore, when the power could not be

ncreased further in these conditions, the CHF was judged as being

etected.

6

When the CHF was detected, the voltage was reduced by 0.25 V

nd the measurement was obtained once steady-state was ver-

fied. The state of the system was considered steady when the

emperature measurements did not change more than 0.15 °C for

0 s. Temperatures, pressures, and mass flow rate were sampled

t 10 Hz, while the power supplies were sampled at 1.5 Hz for

0 s. The infrared camera recorded 30 frames during the logging

fter which the high-speed camera recorded 150 frames at 10 kHz.

he high-speed camera was mainly used to inspect and ensure the

wo-phase flow stability in our recordings, i.e. verify the absence

f appreciable vapor backflow.

.4. Data reduction

The effective geometric dimensions were used for each mi-

rochannel heat sinks throughout the data reduction as reported in

able 2 . The heat loss from the test section was determined from

series of single-phase liquid flow measurements through the test

ection and from a series of measurements with vacuum in the

est section. The data indicated that the refrigerant mass flow had

n insignificant effect on the heat loss compared with the tem-

erature difference between the heater and the ambient. The re-

ression of the data indicated an R

2 value of 89.7% and a standard

eviation of 1.13 W. Moreover, the estimated heat losses were in

etween (12.2–20.2) W for all the measurements. Considering the

eat loss, the footprint heat flux was calculated by

′′ f p =

V · I − ˙ Q e

A f p

(1)

here V was the voltage measured by the power supply, I was the

urrent through the heater, ˙ Q e was the heat loss to the environ-

ent and A f p was the footprint area. The wall heat flux was calcu-

ated based on the effective wetted perimeter P wet , i.e. employing

he fin efficiency η for rectangular fins with insulated fin tip.

′′ w

=

q ′′ fp

(W ch + W fin

)( W ch + 2 ηH ch )

=

q ′′ fp

(W ch + W fin

)P wet

(2)

=

tanh

(m H f in

)(m H f in

) (3)

2 =

h · 2

(W f in + L h

)k Cu

(W f in L h

) (4)

here k Cu was the thermal conductivity of copper and η was the

n efficiency. The heat transfer coefficient was calculated by

=

q ′′ fp

( T w

− T sat ) (5)

here the wall temperature T w

was calculated by using the av-

rage temperature of the heater T h and assuming 1-D conduction

hrough the silicon chip, the solder and the copper.

w

= T h − �T Si − �T Sn − �T Cu

�T i = k i q ′′ fp /t i i ∈ { Si , Sn , Cu } (6)

The temperature of the heater was taken as the average of the

TD temperatures. Moreover, the IR temperature readings needed a

ero calibration by using the RTDs, and therefore the two averages

ere almost identical. The saturation temperature was calculated

y using the average pressure inside the channels. The channel in-

et and outlet pressures were calculated by

p ch,in = p in − �p p,in − �p o − �p c (7)

p ch,out = p out + �p p,out + �p e (8)

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Table 4

Range and uncertainties of reduced variables based on 95% confidence.

Variables Range Relative Absolute

uncertainty, % uncertainty

Mass flux, G , kg/m

2 s 333 - 1260 4.4 - 6.8 14.8 - 85.3

Critical vapor quality, x cr 0.28 - 0.98 0.005 - 0.017

Footprint CHF, q ′′ f p

, W/cm

2 228 - 678 1.0 - 1.8 4.2 - 6.7

Wall CHF, q ′′ w , W/cm

2 49 – 167 3.4 - 5.3 2.1 - 7.2

Saturation temperature, T sat , °C 29.7 - 41.5 0.06 - 0.11

Inlet subcooling, �T sub , K 1.3 - 14.7 0.17 - 0.21

w

a

s

t

t

d

fl

t

t

m

s

t

0

e

x

w

w

e

w

2

t

t

u

T

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a

p

T

a

T

t

p

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m

c

s

h

n

a

e

p

2

o

g

p

T

i

r

e

h

o

t

s

(

w

i

a

T

a

e

t

a

t

3

c

a

r

b

3

r

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T

t

“

w

i

m

f

c

u

s

fl

e

f

f

F

e

f

h

a

t

here the liquid pressure drop through the inlet plenum �p p,in

nd the orifices �p o were obtained from a series of adiabatic pres-

ure drop tests, and where the two-phase pressure drop through

he outlet plenum �p p,out was obtained from a series of adiabatic

wo-phase pressure drop tests using the preheater. The pressure

rop tests were carried out for all fluids and regressed for each

uid individually. The pressure drop of the contraction �p c and

he expansion �p e were calculated similarly to [2] . The saturation

emperature drop inside the channels was (0.1 to 1.9) K for all the

easurements. In this context, it should be mentioned that the

aturation temperature difference between the absolute pressure

ransmitters and the differential pressure transmitter was (0.05–

.1) K.

The critical (outlet) vapor quality was calculated by using the

nergy balance

cr =

q ′′ w

Gh fg

· P wet L h A ch

− �h sub

h fg

(9)

=

q ′′ w

Gh fg

· 4 L h D wet

− �h sub

h fg

(9)

here the definition of the hydraulic diameter using the effective

etted perimeter can be used to invoke the effective wetted diam-

ter D wet . This diameter differs only in the use of the fin efficiency

hen compared with the heated equivalent diameter D he .

.5. Experimental uncertainty

The experimental uncertainties resulting from the data reduc-

ion were calculated by using the single sample error propaga-

ion method by Kline and McClintock (1953). The uncertainty val-

es in Table 3 were used together with the standard deviations in

able 2 with 2 σ coverage factor. Similarly, the standard deviation

f the heat loss regression was used with 2 σ coverage factor. An

dditional uncertainty was estimated to 0.5 V for the power sup-

ly considering the 0.25 V resolution of the CHF detection method.

he uncertainties of the layer thicknesses were estimated to (2, 4.5

nd 20) μm for the silicon, the solder and the copper, respectively.

he liquid pressure drop predictions through the inlet plenum and

he orifices were estimated to have 5% uncertainty while the two-

hase pressure drop predictions through the outlet plenum were

stimated to have 10%. These numbers were chosen by keeping in

ind the use of adiabatic regressions in diabatic measurements.

Table 4 summarizes the absolute and relative single sample un-

ertainties based on 95% confidence interval as well as the mea-

urement range. The uncertainty on the mass flux was relatively

igh due to the effect of dimensional uncertainties in small chan-

els on the cross-sectional area. Similarly, these effects were prop-

gated from the footprint CHF to the channel wall CHF in terms of

ffective heated area. The uncertainties were well accepted for the

urpose of the study.

.6. Experimental design

The nominal experimental design conditions were chosen based

n the second objective of the current study, which was to investi-

7

ate the effects of main parameters on the CHF. The mass flux was

arameterized by 9 values from (325 to 1200) kg/m

2 s nominally.

he inlet subcooling was parameterized by (2, 5, 9 and 12) K nom-

nally. The saturation temperature at the test section outlet was pa-

ameterized by (30 and 40) °C nominally. The effect of the heated

quivalent diameter was parameterized by the two microchannel

eat sink designs. Finally, the experimental design includes the use

f low-GWP refrigerants R1234yf and R1234ze(E) as well as the

raditional R134a. In total 432 data points were obtained. At the

aturation temperature of 30 °C, the highest level of subcooling

12 K) was not always reached for the lower mass fluxes, which

as due to the minimum cooling temperature of the central cool-

ng system in the laboratory (ranging from 16 °C to 18 °C) as well

s the heat uptake through the pump and from the surroundings.

his meant that the highest inlet subcooling was around (8–10) K

t the lowest mass flux and 30 °C saturation. Furthermore, the av-

rage saturation temperature increased with the mass flux due to

he increased pressure drop. All the experimental data points are

vailable in the supplemental material (table with channel values)

o allow for further usage.

. Results

This section presents the results for both the footprint CHF and

hannel wall CHF. The effects of the main parameters are detailed

nd the results are compared with correlations. Finally, a new cor-

elation including the effect of the inlet subcooling is presented

ased on the Katto and Ohno [19] formulation.

.1. Footprint critical heat flux

Figs. 5–7 present the footprint CHF results obtained in the cur-

ent work for each of the fluids investigated. Generally, the results

ndicate that the geometry “F” achieves the highest footprint CHF.

his is also expected since the area enhancement with respect to

he footprint is larger for geometry “F” compared with geometry

D”. The general trends are similar for all geometries and fluids,

hile their values differ.

The increase of the CHF with refrigerant mass flux is confirmed

n all cases. In some cases, the CHF reaches an asymptote as the

ass flux is increased (especially at lower inlet subcooling and

or geometry “D”), while most other cases suggest a further in-

rease in CHF with mass flux beyond the current examined val-

es. The effect of the inlet subcooling is consistent in the mea-

urements at higher mass fluxes. At the highest nominal mass

ux, the footprint CHF increases around (10–20) % from the low-

st to the highest inlet subcooling with more relative significance

or geometry “D” having wider channels. For all the fluids, the

ootprint CHF decreases with increasing saturation temperature.

rom (30 to 40) °C the CHF decreased around 10% at the high-

st mass fluxes. Consequently, the highest values for each fluid are

ound at the low saturation temperature, high inlet subcooling and

igh mass flux. For geometry “F”, these values are (678.5, 526.6

nd 644.0) W/cm

2 for R134a, R1234yf and R1234ze(E), respec-

ively. For geometry “D” these values are (581.5, 440.8 and 528.4)

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 5. Footprint CHF vs. mass flux for R134a. Average saturation temperature was 31.3 °C (a) and 40.6 °C (b). Average subcooling (SC) in K indicated for each curve with

respect to geometry “D” and “F”.

Fig. 6. Footprint CHF vs. mass flux for R1234yf. Average saturation temperature was 30.5 °C (a) and 40.0 °C (b). Average subcooling (SC) in K indicated for each curve with

respect to geometry “D” and “F”.

W

1

r

b

a

c

3

p

C

t

4

t

s

C

t

v

t

m

T

a

e

/cm

2 for R134a, R1234yf and R1234ze(E), respectively, i.e. around

00 W/cm

2 less for geometry “D” for all fluids. Thus, the traditional

efrigerant R134a shows the highest CHF values, closely followed

y refrigerant R1234ze(E), but well beyond R1234yf. These values

re about twice as high as the previous CHF investigations indi-

ated in Table 1 . This is discussed in Section 4.

.2. Channel wall CHF

The effect of the channel size on the CHF is obtained by com-

aring the channel wall CHF. Fig. 8 a presents the channel wall

HF versus the mass flux for R1234ze(E) at 30 °C nominal satura-

ion temperature. Fig. 9 a presents the same results for R1234yf at

8

0 °C nominal saturation temperature. The geometry “D” achieves

he highest wall CHF for all values when compared at similar inlet

ubcooling and similar mass flux. This is consistent with common

HF correlations and observations that do predict an increase of

he wall CHF with heated equivalent diameter.

The wall CHF presented in terms of the corresponding critical

apor quality can be observed in Figs. 8 b and 9 b. By employing

he critical vapor quality, the wall CHF results converge approxi-

ately to the same curves with less variation for both geometries.

his result is suggested to be expressed by the effective wetted di-

meter ( Eq. (9 )), while the remaining differences are due to the

ffect of the inlet subcooling at lower critical qualities (high mass

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 7. Footprint CHF vs. mass flux for R1234ze(E). Average saturation temperature was 31.1 °C (a) and 40.5 °C (b). Average subcooling (SC) in K indicated for each curve

with respect to geometry “D” and “F”.

Fig. 8. Wall CHF vs. mass flux (a) and critical vapor quality (b) for R1234ze(E) at 31.1 °C. Average subcooling (SC) in K indicated for each curve with respect to geometry “D”

and “F”.

fl

a

n

F

f

s

c

p

i

g

r

w

t

h

s

i

3

t

T

t

d

uxes) and probably smaller variations in dry-out or CHF location

t higher critical vapor qualities (low mass fluxes).

Fig. 10 a shows the effect of the inlet subcooling on the chan-

el wall CHF for R1234ze(E) at 30 °C nominally for geometry “D”.

ig. 10 b shows the same results for R1234yf at 40 °C nominally

or geometry “F”. These figures exemplify the ranges of the inlet

ubcooling that were measured in the current work. The figures

learly indicate the effects of the inlet subcooling and that this

arameter is significant for the higher mass flux values while be-

ng insignificant for the lower mass flux values. This result is sug-

ested to be caused by the relative importance of the subcooled

egion and the two-phase region in terms of enthalpy change, for

hich the subcooled region �h becomes larger with respect to

sub e

9

he two-phase region h f g x cr at higher mass fluxes. On the other

and, these results could also be related to the two-phase flow in-

tability, which generally decreases with increasing inlet subcool-

ng [ 2 , 8 ].

.3. Comparison with literature correlations

The CHF experiments are compared with 12 selected correla-

ions from the literature in Fig. 11 , which are sorted alphabetically.

he corresponding prediction statistics are given in Table 5 . The

able indicates the mean absolute error (MAE), ±20% and ±30%

ata point coverages, as well as remarks on channel type(s), heated

quivalent diameters and fluids used to develop the correlations

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 9. Wall CHF vs. mass flux (a) and critical vapor quality (b) for R1234yf at 40.0 °C. Average subcooling (SC) in K indicated for each curve with respect to geometry “D”

and “F”.

Fig. 10. Wall CHF vs. inlet subcooling for R1234ze(E) at 30.8 °C for geometry “D” (a) and R1234yf at 40.0 °C for geometry “F” (b). Average mass flux indicated for each curve.

(

e

T

p

t

t

c

e

e

a

a

d

t

l

s

v

t

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l

b

[

c

[

l

s

w

u

table sorted by channel type). The selected correlations cover gen-

ral conventional channels and general microchannel correlations.

he word “general” is used here for highlighting the use of inde-

endent experimental data from different sources. Those correla-

ions without the word “general” use exclusively experiments from

heir own laboratory. The smallest channel diameters cover the

urrent experiments or are very close for most of the correlations,

xcept for the correlations by Katto and Ohno [19] and Shah [20] .

The best statistics are obtained with the correlations by Bow-

rs and Mudawar [7] (MAE = 8.7%), Wu et al. [21] (MAE = 8.5%)

nd Wojtan et al. [9] (MAE = 16.8%). The former two correlations

re able to capture more than 90% of the current experimental

ata points within 20% error. This is very accurate, especially for

he general microchannel correlation by Wu et al. [21] . The corre-

10

ation by Katto and Ohno [19] provides the fourth-best compari-

on (MAE = 22.1%), which is also very accurate for a general con-

entional channel correlation. These four correlations predict more

han 97.0% of our experimental data points within 30% error. It is

orth highlighting the variability of the predictions by the corre-

ations, which can be seen in Fig. 11 . For example, the correlations

y Ko ̧s ar and Peles [13] , Kandlikar [22] and Kumar and Kadam

23] show a relatively large spread in the predictions, while the

orrelations by Katto and Ohno [19] , Shah [20] and Tibiriçá et al.

24] , indicate a low spread of the predictions, even though the two

atter correlations are well under-predicting our experimental re-

ults. Furthermore, some of the correlations indicate a trend to-

ards over-predicting the current experiments at higher CHF val-

es (Anwar et al. [25] , Mikielewicz et al. [26] ) while others in-

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 11. Parity plot of selected CHF correlations vs. the measured CHF values.

11

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Fig. 11. Continued

12

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Table 5

Prediction statistics of selected CHF correlations.

Correlation MAE [%]

Coverage [%] Remarks:

±20 % ±30 % Channel type D he [mm] Fluids

Anwar et al. [25] 38.1 41.3 50.8 single circular channels 0.64–1.7 refrigerants incl. hydrocarbons

Mikielewicz et al. [26] 63.7 10.4 20.6 single circular channels 1.15, 2.3 R134a, R123, ses 32, ethanol

Wojtan et al. [9] 16.8 63.5 98.4 single circular channels 0.5, 0.8 R134a, R245fa

Ong and Thome [10] 32.4 12.0 32.1 single circular and multi

rectangular channels

0.35–0.88 R134a, R236fa, R245fa

Bowers and Mudawar [7] 8.7 95.8 100 multi rectangular channels 2.54, 0.51 R113

Ko ̧s ar and Peles [13] 63.5 7.2 15.0 multi rectangular channels 0.29 R123

Tibiriçá et al. [24] 59.0 0.0 0.0 general single circular

channels

0.24–6.92 water, refrigerants, nitrogen

Kandlikar [22] 50.7 1.4 10.2 general microchannels 0.1–3 water, refrigerants

Kumar and Kadam [23] 34.9 28.4 40.2 general microchannels 0.237–1.6 R123, R134a, R245fa, R236fa

Wu et al. [21] 8.5 91.5 100 general microchannels 0.223–2.98 water, refrigerants, nitrogen

Katto and Ohno [19] 22.1 38.6 97.0 general conventional channels water, refrigerants, cryogens

Shah [20] 47.0 0.0 0.2 general conventional channels water, refrigerants, cryogens,

chemicals, liquid metals

d

a

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3

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a

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M

i

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q

q

w

u

d

r

m

b

p

9

s

o

m

4

o

t

c

w

Fig. 12. Parity plot of Eq. (10 ) vs. the measured CHF values.

T

s

h

h

h

v

t

f

h

t

a

m

d

6

T

i

f

t

icate a smaller increasing trend (Bowers and Mudawar [7] , Ong

nd Thome [10] , Shah [20] , Tibiriçá et al. [24] , Wojtan et al. [9] ).

n the other hand, the experiments performed by Anwar et al.

25] included high vapor quality measurements ( > 0.8), which cor-

esponds to the low CHF values and indicates good agreement

ere.

.4. Development of a new correlation

Fig. 11 suggests that the variability of the current experimen-

al data can be well predicted using the formulations by Katto

nd Ohno [19] . These were also applied in other correlations (An-

ar et al. [25] , Bowers and Mudawar [7] , Ong and Thome [10] ,

ikielewicz et al. [26] , Wojtan et al. [9] ), however, none of them

ncluded the effect of inlet subcooling in their correlations. This

as due to the absence of an explicit effect of inlet subcooling in

heir data. The current experimental data suggests including the

ffect of subcooling. Thus, all the Katto and Ohno formulations

ere analyzed and the best comparison was found with the fol-

owing formulation:

′′ w

= q ′′ w 0

(1 + 1 . 68

�h sub

h f g

)(10)

′′ w 0 = 0 . 257

(ρg

ρ f

)0 . 17

We L −0 . 27

(L h

D he

)−0 . 36

Gh fg (11)

here the q ′′ w 0

represents the critical heat flux with saturated liq-

id inlet. The correlation can predict the current experimental

ata with 4.0% MAE while having more than 93.3% within 10% er-

or. The regression was performed using the nonlinear regression

odel in Matlab 2018b. All coefficients were found to have proba-

ility values close to numerical precision, which means that all the

arameters are statistically significant. The adjusted R

2 value was

6.5%. On the other hand, using Eq. 11 alone without the effect of

ubcooling resulted in an MAE of 5.5% and an adjusted R

2 value

f 92.9%. Fig. 12 shows the comparison of the correlation vs. the

easurements.

. Discussion

The current target heat flux for electric vehicles is 250 W/cm

2

n chip level. It means that there is a large design space, in which

wo-phase flow boiling with the current heat sinks and refrigerants

ould be used in such application. The maximum footprint CHF

as twice as high as the previous CHF investigations indicated in

13

able 1 and satisfied the first objective of the paper. The main rea-

on is believed to be a combination of the higher surface area en-

ancement and channel cross-sectional area owing to the narrow

igh aspect ratio channels as well as the short heated length. The

ighest channel wall CHF values were also much higher than pre-

ious investigations [ 6 , 13 , 14 ]. The fin efficiency ranged from 0.70

o 0.90 indicating that the fins could have been made even higher

or increasing CHF. Compared with lower aspect ratio channels, the

igh cross-sectional area allows for more refrigerant mass flow at

he same mass flux. It results in much more liquid in the channels

s well as a higher hydraulic diameter and lower pressure drop.

The refrigerant R1234yf did not provide as good CHF perfor-

ance as the refrigerants R134a and R1234ze(E). Averaging all the

ata points for each fluid, indicated a CHF reduction of 22.5% and

.8% comparing R1234yf and R1234ze(E) with R134a, respectively.

he effects of some selected saturation properties can be observed

n Table 6 at 35 °C including the normalization of the low-GWP re-

rigerants with R134a. The changes in the latent heat of vaporiza-

ion are believed to be the main factor dominating the CHF for the

M.R. Kærn, G. Criscuolo, K.E. Meyer et al. International Journal of Heat and Mass Transfer 180 (2021) 121840

Table 6

Selected refrigerant properties at 35 °C and their normalization ( ∗) with respect to R134a.

Refrigerant h f g σ ρ f ρg k f c p, f μ f μg

[kJ/kg] [N/m] [kg/m

3 ] [kg/m

3 ] [kW/m K] [J/kg K] [Pa s] [Pa s]

R134a 168 6.7 � 10 −3 1168 43.4 76.9 1471 172 � 10 −6 12.1

R1234yf 137 5.0 � 10 −3 1054 50.3 60.5 1443 135 � 10 −6 12.0

R1234ze(E) 159 7.6 � 10 −3 1129 35.3 70.9 1422 168 � 10 −6 12.8

R1234yf ∗ 0.81 0.74 0.90 1.16 0.79 0.98 0.79 0.99

R1234ze(E) ∗ 0.95 1.12 0.97 0.81 0.92 0.97 0.98 1.06

d

m

s

t

a

a

f

c

r

m

O

f

v

t

(

R

f

f

u

l

a

t

s

g

i

c

r

b

c

o

S

f

i

h

m

e

t

c

s

o

c

c

t

i

b

i

b

r

f

v

l

o

5

a

o

t

a

R

e

2

g

f

p

l

l

b

o

C

b

i

a

1

D

c

i

C

g

C

e

i

c

s

A

d

0

S

f

2

R

ifferent fluids. The reductions in the latent heat of vaporization

atch closely the above reductions in the averaged CHF values. A

imilar reduction is present for the liquid density and the liquid

hermal conductivity, whereas the liquid-specific heat capacities

re similar. The surface tension and the vapor density are highest

nd lowest, respectively, for refrigerant R1234ze(E), and wise versa

or refrigerant R1234yf. It may suggest that these properties could

ompensate for each other. On one hand, the lower vapor density

esults in higher vapor velocity and interfacial shear forces, pro-

oting interfacial waves and liquid entrainment in the vapor core.

n the other hand, the higher surface tension and liquid cohesive

orces result in less interfacial waves and liquid entrainment in the

apor core.

The power required to pump the fluids through the test sec-

ion ( Fig. 2 ) was very low compared to the CHF, i.e. a maximum of

0.19, 0.25 and 0.21) W were obtained for R1234yf, R1234ze(E) and

134a, respectively. The pumping power was found to be similar

or the refrigerants R1234yf and R134a while being slightly larger

or R1234ze(E). Even though the R1234yf had a fairly lower liq-

id viscosity, it also had a larger liquid volume flow due to the

ower liquid density and thus indicating a similar pumping power

s R134a. The larger pumping power of R1234ze(E) was suggested

o be attributed to the higher vapor viscosity. Moreover, the pres-

ure drop through the channels was below (199 and 336) mbar for

eometry “D” and “F”, respectively.

Both geometries achieved a low average surface roughness dur-

ng manufacturing and the values lie in the lower end of what

ould be achieved with metal milling. These differences in surface

oughness are however expected to play a minor role during flow

oiling in microchannels. Kandlikar and Spiesman [ 30 ] found small

hanges in the onset of nucleate boiling for a relatively large range

f average surface roughness (0.188, 0.363, 0.716 and 3.064 μm).

imilar observations were identified by Jones and Garimella [ 31 ]

or rougher surfaces (1.4 μm, 3.9 μm and 6.7 μm), indicating little

nfluence on the onset of nucleate boiling as well as the critical

eat flux.

The effects of the main parameters were in agreement with

ost other findings, i.e. CHF increases with mass flux and diam-

ter while CHF decreases with saturation temperature. Regarding

he inlet subcooling, it was also found to be consistently signifi-

ant in our results at the higher mass fluxes, while having an in-

ignificant effect at lower mass fluxes. This is different from many

ther investigations [ 6–8 , 10 , 12 , 14 , 27 ], which found the inlet sub-

ooling to be moderate or insignificant. Furthermore, the inlet sub-

ooling was found to be statistically significant in the regression of

he CHF measurements. The effect of subcooling could be dimin-

shed as suggested by Qu and Mudawar [8] , by two-phase insta-

ilities such as vapor backflow to the inlet plenum. In the current

nvestigation, we attempted to reduce the two-phase instabilities

y having both upstream throttle valve, inlet orifices and isolated

efrigerant flow to and from the channels. The two-phase flow was

urther examined with a high-speed camera to ensure insignificant

apor backflow. The effect of short heated lengths is further be-

ieved to improve the stability, by reducing the relative importance

f the orifice and the channel pressure drop.

14

. Conclusion

This paper investigated saturated CHF in narrow and high

spect ratio multi-microchannels. These channels were capable

f achieving footprint CHF up to (678.5 and 581.5) W/cm

2 for

he thinner and wider channels, respectively. Refrigerant R134a

chieved the highest CHF values, closely followed by refrigerant

1234ze(E), indicating a 6.8% lower CHF on average. The refrig-

rant R1234yf achieved considerably lower CHF values, indicating

2.5% lower CHF compared with R134a. These reductions are sug-

ested to be attributed to the lower latent heat of vaporization

or R1234yf and R1234ze(E). The effects of some other important

roperties, such as the vapor density and surface tension, are be-

ieved to compensate for each other.

The effects of the main parameters were consistent with most

iterature. Regarding the effect of inlet subcooling, it was found to

e consistently significant at higher mass fluxes. The applicability

f some general correlations [ 19 , 21 ] was found to provide accurate

HF predictions, while more specific correlations [7] could provide

etter accuracy. A new modified Katto and Ohno correlation that

ncluded the effect of subcooling was finally proposed, indicating

mean average error of 4.0% and 93.3% of the data points within

0% error.

eclaration of Competing Interest

The authors declare that they have no known competing finan-

ial interests or personal relationships that could have appeared to

nfluence the work reported in this paper.

RediT authorship contribution statement

Martin Ryhl Kærn: Conceptualization, Methodology, Investi-

ation, Writing – original draft, Funding acquisition. Gennaro

riscuolo: Conceptualization, Methodology, Writing – review &

diting. Knud Erik Meyer: Conceptualization, Methodology, Writ-

ng – review & editing, Supervision. Wiebke Brix Markussen: Con-

eptualization, Methodology, Writing – review & editing, Supervi-

ion.

cknowledgments

This research was supported by the Danish Council for Indepen-

ent Research – Technology and Production Sciences (DFF – 7017–

0356).

upplementary materials

Supplementary material associated with this article can be

ound, in the online version, at doi: 10.1016/j.ijheatmasstransfer.

021.121840 .

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