Immersion cooling nucleate boiling of high power computer chips

Energy Conversion and Management 53 (2012) 205–218

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Immersion cooling nucleate boiling of high power computer chips Mohamed S. El-Genk ⇑ Institute for Space and Nuclear Power Studies, MSC01 1120, 1 University of New Mexico, Albuquerque, NM 87131-0001, USA Chemical and Nuclear Engineering Department, University of New Mexico, Albuquerque, NM 87131-0001, USA Mechanical Engineering Department, University of New Mexico, Albuquerque, NM 87131-0001, USA

a r t i c l e

i n f o

Article history: Received 27 December 2010 Received in revised form 3 August 2011 Accepted 13 August 2011 Available online 1 October 2011 Keywords: Nucleate boiling Critical heat flux Immersion cooling spreaders Electronics cooling Dielectric liquids High-power computer chips

a b s t r a c t This paper presents experimental results of saturation and subcooled boiling of FC-72 and HFE-7100 dielectric liquids on uniformly heated, 10  10 mm porous graphite (PG) surfaces for potential applications to immersion cooling of high power computer chips. The experiments investigated the effects of surface inclination, from upward-facing (0°) to downward-facing (180°), and liquid subcooling from 0 to 30 K on nucleate boiling heat transfer coefficient and critical heat flux. The presented experimental data and correlations for natural convection of dielectric liquids on PG and plane surfaces are important for cooling chips while in the standby mode when surface heat flux <20 kW/m2. The experimental curves of the nucleate boiling heat transfer coefficient for FC-72 dielectric liquid in the upward-facing orientation are used in 3-D thermal analysis for sizing and quantifying the performance of copper (Cu), PG and PG–Cu composite spreaders for removing the dissipated thermal power by an underlying 10  10 mm computer chip with non-uniform heat dissipation. The 2 mm-thick spreaders are cooled by either saturation or 30 K subcooled nucleate boiling of FC-72 and the composite spreader consists of 0.4 mm-thick surface layer of PG and 1.6 mm-thick Cu substrate. Ó 2011 Elsevier Ltd. All rights reserved.

1. Introduction The ever increasing transistors’ density and processing speed of high power computer chips rapidly increase the thermal power dissipation and junctions’ temperature and result in surface hotspots. These effects are concerns in cooling the chips and computer processing units (CPUs). The local heat flux at the hot-spots is >3 times the chip’s average surface and the induced thermal stresses could decrease the service life and increase the failure frequency of the chip. The time to failure typically decreases exponentially with increased junction’s temperature. In addition, the total heat dissipation rate by a high power chip could reach or exceed 100 W/cm2, requiring effective cooling solutions. Immersion cooling by nucleate boiling of dielectric liquids such as FC-72 and HFE-7100 has been recognized to offer potential advantages for cooling high power chips. These include high heat transfer coefficient, low junctions temperature (<85 °C) and a relatively uniform surface temperature. Dielectric liquids of PF-5060, FC-72 and HFE-7100 (Table 1) are being considered for immersion cooling applications. They are chemically compatible with most structure and heat spreader materials, environmentally friendly and have relatively low saturation ⇑ Address: Institute for Space and Nuclear Power Studies, MSC01 1120, 1 University of New Mexico, Albuquerque, NM 87131-0001, USA. Tel.: +1 505 277 5442; fax: +1 505 277 2814. E-mail address: [email protected] 0196-8904/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.enconman.2011.08.008

temperatures (54–64 °C at atmospheric pressure). A challenge with these liquids, however, is the excursion in the surface temperature prior to initiating nucleate boiling. On plane copper (Cu) surfaces such an excursion could exceed 25 K [1–5]. The low surface tension (Table 1) and highly wetting dielectric liquids (static contact angle on most metal and silica surfaces <5°) flood the surface crevices, delaying bubbles nucleation until reaching high temperatures. Temperature excursion prior to initiating nucleate boiling is undesirable because of concern that junctions’ temperature might exceed the industry recommendations (typically <85 °C and <110 °C in special high temperature applications). Numerous experimental investigations have been reported on enhancing nucleate boiling of dielectric liquids using micro- and macro-structured and porous surfaces; surfaces with reentrant cavities, fabricated pores or micro-fins; roughened and microstructured surfaces; porous graphite, metal foam and surfaces with micro-porous coatings [4–46]. Reported results demonstrated increases in the nucleate boiling heat transfer coefficient, hNB, and the critical heat flux (CHF), lower surface temperatures and a reduction or elimination of the temperature excursion prior to boiling incipience. A direct application of the surfaces for enhancing nucleate boiling in immersion cooling of high power chips is to apply them onto the exposed surface of a metal (Cu or aluminum) heat spreaders for removing the dissipated thermal power by underlying chips. Such spreaders would increase the removal of the dissipated thermal power by the chips, while keeping the junctions’ temperature sufficiently low [32,33].

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Table 1 Properties of dielectric liquids for electronic cooling applications.

a

Saturation properties

HFE-7100a

PF-5060a

FC-72a

Boiling point (°C) Freeze point (°C) Molecular weight (g/mole) Liquid density (kg/m3) Vapor density (kg/m3) Viscosity (kg/ms)  104 Specific heat (J/kg K) Latent heat (kJ/kg) Th. conductivity (W/m K) Surface tension (N/m)

54 135 250 1387 8.1 3.9 1241 113.4 0.063 0.011

51.4 190 338 1606 11.3 4.46 1094 96.4 0.054 0.010368

51.5 90 338 1619 11.4 4.4 1094 96.4 0.054 0.0086

Due to high elevation in Albuquerque, NM (0.085 MPa).

A summary of experimental results of recent investigations of nucleate boiling of FC-72 and HFE-7100 dielectric liquids (Table 1) on porous graphite (PG) by El-Genk and Parker [37] are presented in this paper. These results include the measured pool boiling and the nucleate boiling heat transfer coefficient curves, CHF and natural convection data for these dielectric liquids on PG and plane Cu. The presented results also show the effects of surface inclination (from 0° (upward-facing) to 180° (downward-facing) and liquid subcooling up to 30 K on nucleate boiling heat transfer coefficient and CHF. The obtained experimental curves of the nucleate boiling heat transfer coefficient of FC-72 dielectric liquid in the upward-facing orientation [13,37] are used in 3-D thermal analysis for sizing and quantifying the performance of copper (Cu), PG and PG–Cu composite spreaders for removing the dissipated thermal power by an underlying 10  10 mm computer chip with non-uniform heat dissipation [32,33]. The spreaders are 2 mm-thick and the PG–Cu composite spreader consists of 0.4 mm-thick surface layer of PG and 1.6 mm-thick Cu substrate. The composite spreader takes advantage of the nucleate boiling enhancement on PG and the good heat spreading property of Cu. The dissipated power by the underlying chip is removed from the exposed surface of the spreader by nucleate boiling of saturated or 30 K subcooled FC-72 liquid. The next section briefly describes the test section and the pool boiling experiments setup and procedures.

2. Experiments setup and procedures The test section and the experimental facility for conducting the boiling experiments have been described in details elsewhere [1,2,13,23,36,37] and only briefly summarized herein. Fig. 1 presents Scanning Electron Microscope (SEM) images of the porous graphite surface used in the pool boiling experiments with degassed FC-72 and HFE-7200 dielectric liquids [13,23,36,37]. The 10  10 mm PG block is 3 mm thick. The porous graphite has a highly porous structure with randomly sized interconnected pores and cavities with sizes ranging from <1 lm to tens and hundreds of microns (Fig. 1). This highly porous material (61% volume porosity) has 95% open and interconnected pores and the re-entrant cavities have non-circular openings and ragged non-smooth interior and could be several hundred microns deep (Fig. 1). The porous graphite has anisotropic thermal conductivity, 245 W/m K out of plane and 70 W/m K in plane and is stronger than graphite foams and commercially available with consistent properties. The air solubility in dielectric liquids (Table 1) is quite high (>40%). Thus, for consistency of results, pool boiling experiments are conducted with degassed liquids [13,23,36,37]. It typically takes a few hours of continuous boiling to outgas the dielectric liquid pool in the test vessel prior to conducting the experiments. However, air remains entrapped in the tiny pores and re-entrant cavities of the

porous graphite in the test section. The magnetic stirrer at the bottom of the test vessel speeds up the out-gassing of the liquid and ensures uniform pool temperature prior to conducting the experiments. The experiments setup consists of a sealed test vessel submerged in a polycarbonate hot water bath with electrical heater for minimizing the side heat losses from the vessel. The test vessel has a tightly sealed cover to minimize loss of the volatile dielectric liquids, thus maintaining a constant liquid level above the test section (8 cm) in the experiments. The hot water bath and the two submerged cooling coils in the test vessel maintain the liquid pool temperature constant within ±0.5 K of the saturation temperature in Albuquerque, New Mexico (Table 1) or the desired subcooling in the experiments. A water-cooled copper coil is located below the cover plate of the test vessel to condense the vapor generated in the experiments. The liquid pool temperature in the test vessel, monitored using four submerged K-type thermocouples, is taken as the average of those indicated by the two thermocouples placed a few millimeters from PG surface in the liquid pool. The assembled test section measures 30 mm  30 mm in outside dimensions and is made of a Teflon block with a 1.0 mm deep square cavity (10  10 mm) at the center of the top surface in which a high flux heating element and PG block are placed. The PG block is mounted on top of the heating element using thermally conductive epoxy. It has two 0.6 mm diameter horizontal holes drilled equidistance from the top and bottom surfaces on one side and half way into the block. The measuring tips of the two K-type thermocouples are securely attached to the inside of the holes using a thermally conductive epoxy. The average reading of these thermocouples is taken as the surface temperature for the purpose of constructing the pool boiling curves, after accounting for the temperature drop by conduction to the surface of PG block (61.0 K). The Teflon block is encased in a Lexan frame with closed bottom. The shallow cavity on top of the Teflon block, surrounding the mounted PG block, is filled with translucent epoxy adhesive; flush with the PG surface. The purpose is to prevent forming micro-cracks at the edges of the PG block that would stimulate bubbles nucleation and skew the pool boiling curve. The heat losses through the sides, top and bottom of the assembled test section are calculated using the ANSYS finite element commercial software to be negligibly small [1,2,13,23,36,37]. Thus, the removed power from the surface of the uniformly heated, 10  10 mm PG block equals that generated by the underlying heating element in the test section. This power, determined from the measured voltage across and the electric current provided to the heating element by a DC power supply, is that dissipated from the PG surface for the purpose of constructing the boiling curves. In the experiments, the electrical power supplied to the heating element is increased incrementally by raising the applied voltage by <0.02 V at a time. The magnitude of the voltage increment is decreased gradually as the boiling curve approaches CHF. Following

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Fig. 1. SEM images of the porous graphite surface at different magnifications [13,23,37,47].

each incremental increase in the applied voltage, the surface heat flux and average temperature are recorded only after reaching steady state. This is when the difference between two successive measurements of the surface average temperature is within ±0.2 K. Each of these surface temperatures is the average of 30 readings of each of the two thermocouples embedded into the PG block in the test section. When either of these thermocouples detects an increase in excess of 30 K in two consecutive steady state measurements, it is considered an indication of reaching CHF, and the experiment is terminated. These procedures protect the heating element in the test section from burning out when reaching CHF. The estimated uncertainties in the experiments are ±0.7 K and ±0.10 mV in temperature and applied voltage measurements, ±2% and ±3.9% in nucleate boiling heat flux and heat transfer coefficient, ±3% in CHF and 3–6 K in the surface temperature corresponding to CHF. These uncertainties are based on the methodology outlined by Kline [47]. 3. Results and discussion This section presents and discusses the nucleate boiling and natural convection data of FC-72 and HFE-7100 dielectric liquids on PG [13,23,36,37]. A few photographs of nucleate boiling at the different surface orientations are also presented to illustrate the processes of vapor bubbles nucleation, growth, mixing in the boundary layer. 3.1. Pool boiling of dielectric liquids In boiling heat transfer, while CHF is a practical limit not to exceed, desirable operating points for electronics cooling applications are far enough from CHF near the maximum nucleate boiling heat transfer coefficient, hMNB. This coefficient occurs near the end of

NC: Natural Convection

Region I: Discrete Bubble nucleate boiling Region II: Fully developed nucleate boiling Region III: Bubbles coalescence nucleate boiling

Temp. Overshoot

CHF: Critical Heat Flux Fig. 2. A schematic of typical pool boiling curve of dielectric liquids.

the fully developed nucleate boiling region (region II in Fig. 2). It is much higher than that at CHF and occurs at a lower surface superheat or temperature. Fig. 2 presents a schematic of a typical pool boiling curve of dielectric liquids. It consists of four successive heat transfer regions: (a) natural convection (NC) up to the incipience of nucleate boiling, (b) the discrete-bubbles nucleate boiling (region I) at low heat flux and relatively low surface superheats; (c) the fully-developed nucleate boiling (region II) at higher heat flux and surface superheat than in region I and (d) the bubbles coalescence nucleate boiling (region III), at higher heat flux and surface superheat than in region II, ending with CHF. In the discrete-bubbles region (I), the nucleate boiling heat transfer coefficient, hNB, increases with increased surface temperature, as more bubble nucleation sites become active. In the fully

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5

Reentrant Cavities, FC-72 [34] Reentrant Cavities, FC-72 [34] Micro-finned, FC-72 [24] MPC, FC-72 [35] MPC, FC-72 [32] MPC, FC-72 [35]) + 10% PG, HFE-7100 [23]

4

qNC (W/cm2)

developed nucleate boiling region (II), the density of the active nucleation sites as well as the heat transfer coefficient are highest, as indicated by the steep slope of the pool boiling curve in this region (Fig. 1). The maximum nucleate boiling heat transfer coefficient, hMNB, occurs near the end of region (II) before transitioning to region (III). In region (III), hNB decreases with increased heat flux because of the larger increase in the surface temperature. This temperature increase is caused by the added heat transfer resistance due to the coalescence of growing and rising bubbles at and near the heated surface. This explains why at CHF the nucleate boiling heat transfer coefficient is lower and the corresponding surface temperature, TW, or superheat, DTsat = (Tw Tsat), where Tsat is liquid saturation temperature, are higher than at hMNB. For electronics cooling applications, it is preferable to operate in region II near hMNB, on the left side of the pool boiling curve, where the surface superheat is sufficiently lower than at CHF (Fig. 2). The reported enhancements in nucleate boiling of dielectric liquids on micro- and macro-structured or porous surfaces; surfaces with reentrant cavities, fabricated pores or micro-fins; micro-structured surfaces; porous graphite, and surfaces with micro-porous coatings and metal foam [4–46], are primarily due to the increases in the density of the active sites for bubbles nucleation. The resulting increase in the nucleate boiling heat transfer rate shifts the pool boiling curves to the left, to lower surface superheats. The next section presents natural convection data and correlation for dielectric liquids on porous graphite and micro-porous and plane Cu and silicon surfaces [1,4,13,24,32,34,35].

3

- 10%

PG, FC-72 [13] 2 1.2

qNC = 0.0353 ΔTp

1

(a) Porous and Structured Surfaces 0

0

10

20

30

40

50

5 Cu, FC-72 [35] Cu, FC-72 [4] Cu,FC-87 [32] Cu,FC-72 [32] Cu, FC-72 [35] Cu, HFE-7100 [1] Cu,HFE-7100 [23] Cu, FC-72 [13]

4

qNC (W/cm2)

208

3

+ 7%

- 7%

2 1.2

qNC = 0.0314 ΔTp

1 (b) Plane Surfaces

0

0

10

20

30

40

50

60

ΔTp (K)

3.2. Natural convection of dielectric liquids Computer chips and CPUs typically operate in the standby mode for extended periods of time. In this mode, the dissipation heat flux is typically <20 kW/m2. The experimental data presented in Fig. 3a and b are from investigations performed in our laboratory and reported by others [1,4,13,24,32,34,35]. These figures plot the natural convection heat flux, qNC, versus the surface superheat, DTp, where DTp = (Tw Tp) and Tp is the liquid pool temperature. Fig. 3a presents the natural convection data for different dielectric liquids on porous, micro-porous and structured surfaces and Fig. 3b presents similar data on plane surfaces, for comparison. These figures indicate that for the same DTp, qNC for both FC-72 and HFE-7100 liquids on PG, micro-porous and structured surfaces, is 12.4% higher than on plane surfaces. The obtained and reported natural convection heat flux data for these uniformly surfaces are correlated in terms of surface superheat, DT 1:2 p (Fig. 3a and b). Following boiling incipience, the nucleate boiling heat transfer rate increases with increased heat flux (Fig. 2). Some of the obtained results in our laboratory on enhancements of saturation and subcooled nucleate boiling of FC-72 liquid on PG are presented next, along with the boiling curve of the same liquid on plane Cu for comparison [1,2,13,36,37]. 3.3. Nucleate boiling on porous graphite The experimental results presented in Figs. 4 and 5 are for saturation and subcooled boiling of FC-72 dielectric liquid on plane Cu and PG surfaces in the upward-facing orientation. Figs. 4a and 5a show the pool boiling curves at saturation and 10, 20, and 30 K subcooling and Figs. 4b and 5b present the corresponding curves of the nucleate boiling heat transfer coefficient, hNB. Figs. 4a and 5a plot the nucleate boiling heat flux, q, versus DTp and Figs. 4b and 5b compare the corresponding hNB curves. The last data points on the boiling curves in Figs. 4a and 5a indicate CHF. The pool boiling curves of FC-72 and HFE-7100 liquids on porous graphite showed no or little (<5 K) excursion in surface temperature before

Fig. 3. Natural convection data and correlations of dielectric liquids of porous, micro-porous, structured and plane surface.

initiating nucleate boiling, which occurs at surface superheats as low as 0.5 K. The maximum nucleate boiling heat transfer coefficient, hMNB, is indicated by the solid circle symbols in Figs. 4 and 5. It typically occurs near the end of fully developed nucleate boiling region (region II in Figs. 2, 4a and 5a). On both plane Cu and PG surfaces (Figs. 4b and 5b), hMNB decreases and the corresponding surface temperature increases with increased liquid subcooling. The values of hMNB for FC-72 on plane Cu are much lower than on PG. For saturation boiling of FC-72 on Cu, hMNB is only 8.4 kW/m2 K compared to 34 kW/m2 K on PG and the corresponding surface temperature is 5–10 K lower (Figs. 4b and 5b). The nucleate boiling heat transfer coefficient at CHF is not only lower than hMNB, but occurs at a higher surface superheat (Figs. 4b and 5b). These figures show that the values of hMNB for FC-72 on PG (13–33 kW/m2 K) are >2–3 times those on plane Cu (5.3–8.8 kW/m2 K) (Figs. 4b and 5b). In the pool boiling experiments, the large pores in the PG surface could be partially flooded with the liquid, allowing nucleation of vapor bubbles on the interior surfaces, thus increasing the effective area for heat transfer. This may explain the measured increases in the rate of heat removal by nucleate boiling and CHF on PG; the lower surface temperature increases hNB. The evaporation of the thin liquid films on the interior walls of the large pores in the PG surface may also explain the low surface temperature excursion prior to boiling incipience (<5 K). In contrast, the measured excursions in surface temperature prior to the incipience of nucleate boiling of dielectric liquids on Cu varied from 12 to 25 K. 3.4. Effect of surface inclination on CHF For flexibility in electronics packaging, involving immersion cooling nucleate boiling of dielectric liquids, it is desirable to determine and quantify the effect of changing the surface inclination on

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30

Nucleate boiling of FC-72 Liquid on plane Cu, θ = 0 (upward-facing) ΔTSUB = 0.0 K (sat) (III) ΔTSUB = 10.0 K ΔTSUB = 20.0 K (III) ΔTSUB = 30.0 K hMNB

2

q (W/cm )

20

(III)

(III)

(II)

(II) (II)

10

(II) (I)

(I)

(I)

(a)

(I) 0 0

5

10

15

25

30

35

40

45

50

55

(II)

0.8

(I)

(III) (II)

0.6

(III)

(I)

2

hNB (W/cm K)

20

(I)

(II)

(III)

0.4

(II)

(I)

(III)

0.2

(b) 0

0

5

10

15

20

25

30

35

40

45

50

55

ΔTp (K) Fig. 4. Effect of liquid subcooling on nucleate boiling of FC-72 on plane copper [37].

60

Nucleate Boiling of FC-72 on Porous Graphite, o θ = 0 (upward-facing)

q (W/cm2)

50 40

(III)

(III)

(III)

(II)

30

(III) 20 10 0

(II)

(II)

(II) (I) 0

5

10

15

(I)

(I)

(I) 20

25

30

(a) 35

40

45

50

3.5

hNB (W/cm2K)

ΔTsub ΔTsub ΔTsub ΔTsub

(III)

3.0 (II) 2.5

hMNB

(I)

2.0

= 0 K (sat.) = 10 K = 20 K = 30 K

(III)

1.5

(II) (II)

1.0 0.5

(III)

(I)

(I)

(II) (III)

(b)

(I)

0 0

5

10

15

20

25

30

35

40

45

50

ΔTp (K) Fig. 5. Effect of liquid subcooling on nucleate boiling of FC-72 on porous graphite [37].

the nucleate boiling heat transfer coefficient and CHF. This section presents some results of the pool boiling experiments carried out in our laboratory of saturation and subcooled nucleate boiling of FC-72 and HFE-7100 liquids on PG at different surface orientations from 0° (upward-facing) to 180° (downward-facing) [13,23,37]. Fig. 6a and b compares the obtained results with those reported by other investigators on other surfaces [2,11,39,41–44]. The results in Fig. 6a and b shows that the highest saturation CHF values for HFE-7100 and FC-72 are those measured on PG [11,23], followed by those reported on micro-porous coatings [39] and then on Cu [2,11,23,39,41–44]. In the upward-facing orientation (h = 0°), saturation CHF of FC-72 on PG (30 W/cm2) [13] is 12% higher than that reported on micro-porous coating (26.8 W/cm2) [39] and up to 57% higher than that on plane Cu (19.1 W/cm2) (Fig. 6a). The CHF values of FC-72 on plane Cu are lower than those of HFE-7100 (Fig. 6b) and range from 16.1 W/cm2 to 22.4 W/cm2. Fig. 6a and b shows a large scattering in the reported CHF values by different investigators for saturation boiling of FC-72 and HFE-7100 on plane Cu at the various surface orientations, including 180° (downward-facing). In this orientation, the reported CHF values on plane Cu are 1.6 W/cm2 [39], 2.5 W/cm2 [13] and as much as 4.9 W/cm2 [39]. Such variations result from the uncertainties in the adjustment of the surface in the downward-facing orientation and the procedures used by different investigators for determining CHF in the experiments. The results delineated in Fig. 6a and b shows CHF decreases slowly with increased surface inclination up to 90° then decreases much faster with increased inclination to its lowest value at 180°. The rates of decrease in saturation CHF of FC-72 and of HFE-7100 with inclination up to 90° are similar on porous graphite,

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35

2.2

(a) Saturaion Boiling FC-72 30

2.0

25

1.8

HFE-7100 [11]

CHFR

CHF (W/cm2)

+4%

20 15

10 x 10 mm Cu [23] 10 x 10 mm Cu [39] 12.7 x 12.7 mm Cu [41] 12.7 x 12.7 mm Cu [42] 12.7 x 12.7 mm Cu [43] 10 x 10 mm Micro-Porous Coating [39] 30 mm dia. Cu [44] 10 x 10 mm Porous Graphite [23]

10 5

1.6 1.4

CHFR =1+0.036 ΔTsub

1.2 1.0

(a) Porous Graphite

2.0

Cu, HFE-7100 [2] Cu, HFE-7100 [5] Si, HFE-7100 [46] Cu, HFE-7100 [11]

0 1.8

40

(b) Saturation Boiling HFE-7100 35

CHFR

1.6

30

CHF (W/cm2)

-4%

+ 8%

- 8%

1.4

25 1.2

20

1.0

15

(b) Plane Surfaces 0

10 10 x 10 mm Porous Graphite [11, 23] 10 x 10 mm Cu [2] 30 mm dia Cu [44] 10 x 10 mm Cu [11]

5 0

CHFR =1+ 0.024 ΔTsub

0

30

60

90

120

5

10

15

20

25

30

Liquid Subcooling, ΔTsub (K) 150

180

Fig. 7. Effect of liquid subcooling on CHF of HFE-7100 dielectric liquid on porous graphite.

Inclination, θ (°) Fig. 6. Effect of surface inclination on saturation CHF of HFE-7100 and FC-72 dielectric liquids on various surfaces.

3.5. Effect of liquid subcooling on CHF on PG Subcooled boiling reduces the size of departing vapor bubbles from the heated surface and as they rise through the liquid pool. This limits the accumulation and growth of the bubbles and reduces their coalescence at and near the heated surface. The net effect is increasing CHF and the corresponding surface temperature or superheat. The resulting increase in the surface temperature with increased liquid subcooling decreases the nucleate boiling heat transfer coefficient (Figs. 4b and 5b). The data presented in Fig. 7a and b shows that CHF values of dielectric liquids on various surfaces increase linearly with increased liquid subcooling. However, the value of the subcooling coefficient, CCHF,sub, depends on the surface characteristics and the thermophysical properties of the boiling liquid. For HFE-7100 on PG, CCHF,sub = 0.036/K (Fig. 7a), compared to only 0.024/K on smooth surfaces (Fig. 7b). These results show that CHF of HFE-7100 on PG in the upward-facing orientation increases 3.6% per degree increase in

2.5

CHFsub/CHFsat

micro-porous coating, and plane Cu. However, beyond 90° inclination the rate of decrease in CHF on PG with increased surface inclination is much smaller than reported on micro-porous coatings. Also, CHF values on PG are consistently higher. In the 180° orientation (downward-facing), the saturation boiling CHF of FC-72 on PG of 16 W/cm2 is 53% of that measured in the upward-facing orientation (0°), while that reported on micro-porous coatings [39] of 4.9 W/cm2 is only 18% of that reported for the upward-facing orientation (Fig. 6a). Similar results are presented for HFE-7100 dielectric liquid in Fig. 6b. The results of the effect of liquid subcooling on CHF values on PG and smooth surfaces in the upward-facing orientation are presented and discussed next.

3.0

2.0

CCHF,sub = 0.044, MPC Gassed FC-72 [40]

CCHF,sub = 0.036, PG Degassed HFE-7100 [11,23]

CCHF,sub = 0.049., MPC Degassed FC-72 [40]

1.5

1.0

CCHF,sub= 0.032, PG Degassed FC-72 [11] CCHF,sub = 0.022, SS Gassed and Degasse CCHF,sub = 0.018, RES,Gassed PF-5060 [24,45] & Degassed FC-72 [4,13]

ΔTsub (K) Fig. 8. Comparison of the effect of subcooling on CHF of different gassed and degassed dielectric liquids on different surfaces in the upward-facing orientation [40].

liquid subcooling, compared to 2.4% on plane surfaces. The values of CHF for HFE-7100 dielectric liquid are typically 10–15% higher than of those of FC-72, however, those of the hMNB for FC-72 are 57% higher [23,37]. The value of CHF for FC-72 on PG in the upward-facing orientation increases linearly with subcooling at a rate, CCHF,sub, of 0.032 K 1 (Fig. 8). In this figure, MPC, PG, RES and SS stand for micro-porous coating, porous graphite, roughened and etched surfaces, and structured surfaces, respectively. The highest values of the subcooling coefficient for FC-72 dielectric liquid are those on micro-porous coating (0.044 and 0.049), followed by that on PG (0.032), then those on structured (0.022) and roughened and etched surfaces (0.018) (Fig. 8). The following section presents still photographs of saturation boiling of HFE-7100 dielectric liquid on PG at different surface orientations.

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Fig. 9. Photographs of saturation nucleate boiling of HFE-7100 liquid on PG in the upward-facing orientation.

3.6. Photographs of nucleate boiling on PG Fig. 9 shows photographs of saturation nucleate boiling of HFE7100 dielectric liquid on PG in the upward-facing orientation: (a) in the discrete bubbles region (region I in Fig. 2), near nucleate boiling incipience (Figs. 2 and 9a) and (b) in the fully developed nucleate boiling (region II in Fig. 2), near the transition from the region I (Fig. 9b). The departing single bubbles from the PG surface (Fig. 9a) is about 0.5–0.55 mm in diameter. Such small bubble diameter, compared to 2.5–3 mm for water at atmospheric pressure, is because of the much lower surface tension of HFE-7100 dielectric liquid (Table 1). The bubbles nucleation in Fig. 9a occurs at a few isolated sites on the PG surface. Increasing the surface heat flux increases the number of active site for bubbles nucleation and the population of the growing bubbles on the surface. It also increases the coalescence of the departing and growing bubbles at and near the surface (Fig. 9b). The heat flux in Fig. 9b is 30% of CHF (Fig. 6b). In addition to the increased density of active sites for bubble nucleation on the PG surface, the relatively small departure bubble diameter contributes to the measured enhancements in nucleate boiling of dielectric liquids. Fig. 10 presents photographs of saturation nucleate boiling of HFE-7100 on PG in the vertical orientation at different heat fluxes and surface superheats, DTsat. The images in Fig. 10 are magnified 300%. They show that the two-phase boundary layer at the heated PG surface grows in thickness with increasing heat flux, as the number of vapor bubbles growing at and departing from the surface increases. The thickness of the boundary layer also increases with axial distance along the heated surface as the bubbles detaching at the lower locations rise up by buoyancy and are captured into the boundary layer. The vapor globules are eventually released from the top edge of the PG surface into the liquid pool. While the induced mixing in the boundary layer by the rising and detaching vapor bubbles increases the nucleate boiling heat transfer rate, the added resistance due to vapor accumulation in the boundary layer increases the surface temperature. This in turn decreases the nucleate boiling heat transfer coefficient and CHF (Fig. 6a and b) and increases the boiling thermal resistance. In the downward-facing orientation (h = 180°), the nucleate boiling heat transfer rate is influenced by gravity in a different way than on inclined PG surfaces (0° < h < 180°). In the downward-facing orientation, the inertia of the rising liquid toward the heated PG

surface, to replace the vapor released from the edges of the text section, causes the growing vapor mass and bubbles at the surface to become increasingly flat and spread sideway (Fig. 11). They are typically separated from the heated surface by a thin liquid film (a few microns thick). The high thermal conductance of this thin liquid film transports the heat efficiently from the PG surface to the liquid– vapor interface of the elongated bubbles residing near the heated surface. These bubbles continue to spread laterally and swell downward, perpendicular to the surface. The internal pressure in the residing vapor mass and bubbles near the surface increases with time due to the accumulation of vapor generated at the liquid–vapor interface facing the heated PG surface. The growth of the vapor globules and the sweeping of these globules from the surface toward the edges of the test section into the liquid pool occur cyclically with a certain frequency (Fig. 11). This frequency increases only slightly as the heat flux increases. Based on visual observation in the experiments and the examination of the recorded videos of the boiling surface, the frequency of the formation and growth of the vapor globules and their release from the edges of the PG surface while in the downward facing orientation is 13–15 Hz [37]. The obtained pool boiling and nucleate boiling heat transfer coefficient curves for saturation boiling of FC-72 on PG and plane Cu surfaces (Figs. 4 and 5) are incorporated in the 3-D thermal analysis of the spreaders. The analysis is performed to size and quantify the performance of spreaders for immersion cooling of a high power computer chip with non-uniform heat dissipation. The spreaders investigated are of the same thickness (2 mm) and include Cu, PG and composite PG–Cu. The approach and results of the thermal analysis of these spreaders are presented and discussed next.

4. Spreaders for immersion cooling of a chip with non-uniform heat dissipation This section qualities the potential of immersion cooling nucleate boiling of high-power computer chips using Cu, PG spreaders and a composite spreader comprised of a 0.4 mm-thick surface layer of PG (tPG) and a 1.6 mm thick Cu substrate (tCu). These spreaders have the same thickness (tSP = 2 mm) and are cooled by either saturation or 30 K subcooled nucleate boiling of FC-72 dielectric liquid (Fig. 12). The chip and the overlying spreader are

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Fig. 10. Photographs of saturation boiling of HFE-7100 liquid on PG surface in the vertical orientation.

squares and the 10  10 mm chip with a non-uniform surface heat dissipation (Fig. 13) is centered below the spreader. The composite spreader takes advantage of the enhancements in nucleate boiling of FC-72 dielectric liquid on PG and the good heat spreading properties of Cu. The square footprint areas of the spreaders are determined in the performed 3-D thermal calculations, subject to the following restrictions on the heat removal by nucleate boiling from the exposed surface of the spreaders [32,33]: (a) hNB at the center of the spreaders is 0.90 of that at CHF in the pool boiling experiment with saturation or 30 K subcooled boiling of FC-72 dielectric liquid and (b) hNB at the corners of the spreader equals to that measured in the experiments at 1 K higher surface temperature than at incipient boiling (Fig. 14). The values of hNB along the spreader surface along the surface depend on the local surface temperature (Fig. 14). In the spreaders’ thermal analysis, the interfacial resistances between the underlying chip and the spreaders is neglected, however, the low saturation temperature of FC-72 (Table 1) and the calculated chip maximum surface temperatures would easily accommodate the temperature drop at the interface without causing the junctions temperature to exceed 85 °C. The dissipated thermal power by the underlying chip is assumed non-uniform and has a cosine-like distribution with a peak-to-average heat flux ratio, Umax (Fig. 13). This ratio varies from unity (uniform) to 2.467. The corresponding ratios of the local heat flux at the edge of the chip to the peak heat flux at the chip center, nq, varies from zero when Umax = 2.467 (Fig. 13b) to unity for a uniform surface heat flux, Umax = 1.0 [33]. The effect of the non-uniformity of the dissipated heat at the surface of the chip on the performance of the Cu, PG and the PG–Cu composite spreaders are quantified and compared. The dissipated thermal power by the underlying chip is removed from the surfaces of the spreaders using either saturation or 30 K subcooled nucleate boiling of FC-72 liquid (Figs. 12 and 14). The values of the saturation nucleate boiling heat transfer coefficient, hNB, of FC-72 on PG are much higher than in subcooled boiling

(Figs. 4, 5 and 14). The value of hNB for saturation boiling increases from 20,000 W/m2 K at a surface temperature of 58.5 °C to a peak of 33,000 W/m2 K, then drops linearly with increasing surface temperature. The peak values of the subcooled boiling hNB are much lower and occur at higher surface temperatures than in saturation boiling (Fig. 14a and b). They are 16,800, 13,500, and 12,000 W/ m2 K and occur at surface temperatures of 62, 64.4, and 65.5 °C, when liquid subcooling, DTsub, is 10, 20, and 30 K. The vertical lines in Fig. 14b indicate the values of hNB corresponding to 90% of those at CHF. Results of the 3-D thermal analysis for sizing the surface areas and quantifying the performance of the spreaders are presented in Figs. 15 and 16. They indicate that increasing the non-uniformity of the dissipated heat at the chip surface, or Umax, decreases the total thermal power dissipated by the chip and removed from the spreader surface, but increases the maximum surface temperature at the center of the chip. When cooled with saturation boiling, the PG–Cu composite spreader (DPG = 0.2) dissipates slightly less thermal power than the Cu spreader of the same thickness (2 mm), but has much smaller surface area (Figs. 15 and 16). As the results in these figures show, the total thermal resistance and chip surface maximum temperature with the composite spreader are much lower than with the Cu spreader (Figs. 15 and 16). When cooled with 30 K subcooled nucleate boiling of FC-72, the removed thermal power from the composite spreader (DPG = 0.2) surface is higher, but its surface area, total thermal resistance, and the chip maximum temperature are lower than with the Cu spreader (Figs. 15–17). The PG spreader removes the least amount of thermal power and has the smallest surface area and the highest total thermal resistance (Fig. 16). With 30 K subcooled boiling and Umax = 2.467, the composite spreader removes 9% more thermal power than the Cu spreader and 18% less power than when the heat flux at the chip surface is uniform (Umax = 1.0). The results show that while cooling the spreaders with subcooled nucleate boiling increases the thermal power removed (Fig. 15), it decreases the surface areas of the speeders and increases

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Fig. 11. Photographs of saturation boiling of HFE-7100 liquid on PG in the downward-facing orientation.

both the total thermal resistances and the maximum surface temperature of the chip (Figs. 15 and 16). When Umax = 2.467, the thermal power removed by the composite spreader (DPG = 0.2) cooled by 30 K subcooled nucleate boiling of FC-72 is 72 W and its surface area measures 11  11 mm (Figs. 15 and 16). The removed thermal power decreases to 39.48 W and the surface area of the composite spreader increases to 13.05  13.05 mm, when cooled by saturation nucleate boiling of FC-72. Decreasing the chip’s heat flux peaking ratio, Umax, increases the total thermal power removed from the surface of the spreaders (Figs. 15 and 16). When Umax = 1.175 and the spreader surface is cooled by 30 K subcooled and saturation nucleate boiling of FC-72,

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the removed thermal power by the composite spreader is 94.7 W and 51.13 W (Fig. 15), and the surface area of this spreader measures 12.07  12.07 mm and 14.05  14.05 mm, respectively. As indicated in Fig. 15, the Cu and composite spreaders cooled by saturation nucleate boiling of FC-72 remove almost the same thermal powers, but the maximum chip surface temperature with the former (71–72.4 K) is much higher than with the latter (62.37–63.4 K). When these spreaders are cooled by 30 K subcooled nucleate boiling of FC-72, the removed thermal powers, the chip maximum surface temperatures, and the total thermal resistances increase, but the surface areas of the spreaders decrease (Figs. 15 and 16). The maximum surface temperature of the chip with a PG spreader is slightly higher than with a composite spreader and the dissipated thermal power is significantly lower (Figs. 15 and 16). Increasing the chip’s heat flux peaking ratio, Umax, increases the maximum surface temperature of the chip by less than a couple of degrees, but decreases the thermal power removed from the surface of the spreader by up to 30 W, depending on the spreader type, Cu, PG or composite (Fig. 15). Fig. 16 plots the calculated surface area of the spreaders versus their total thermal resistance. This resistance is the sum of the boiling and the spreading resistances. The composite spreader cooled by saturation nucleate boiling of FC-72 has the lowest total thermal resistance (0.285 °C/W). It increases to 0.68 °C/W, when the composite spreader is cooled with 30 K subcooled nucleate boiling of FC-72 dielectric liquid. The surface area of the composite spreader cooled by saturation nucleate boiling of FC-72 varies from 6.8 to 8.07 cm2 and from 4.93 to 6.0 cm2 when cooled by 30 K subcooled boiling of FC-72 liquid. The lowest surface areas and the highest total thermal resistances are associated with the largest heat flux peaking ratio, Umax = 2.467. Conversely, the spreaders’ largest surface areas and smallest total thermal resistances are when the power dissipation from the chip surface is uniform (Umax = 1.0). The dependence of the spreaders’ performance on the heat flux peaking ratio, Umax, at the surface of the underlying chip is also true for both the Cu and PG spreaders (Fig. 16), but the calculated surface areas and the total thermal resistance for these spreaders are different. The largest thermal resistance is that with a PG spreader cooled by 30 K subcooled nucleate boiling. This spreader has the smallest surface area. The Cu spreader has the largest surface area and the total thermal resistance is higher than with the composite spreader, but it is much lower than with the PG spreader, for the same conditions. When the heat flux peaking ratio Umax = 2.467, the surface area of the composite spreader cooled by saturation and 30 K subcooled nucleate boiling of FC-72 liquid is 6.82 and 4.9 cm2 and the total thermal resistances are 0.284 and 0.68 °C/W. These values are smaller than those with the Cu spreader (12.26 and 11.92 cm2 and 0.51 and 0.83 °C/W) (Figs. 16– 18).

Fig. 12. A PG–Cu composite spreader for removing dissipated thermal power by underlying 10  10 mm chip using nucleate boiling of dielectric liquid.

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Fig. 13. Ratios of the peak and edge heat fluxes for the underlying chip with non-uniform heat dissipation [33].

With a PG–Cu composite spreader cooled by saturation and 30 K subcooled nucleate boiling of FC-72 liquid, the chip’s maximum surface temperatures of 62.37 K and 72.2 K are much lower than those with the Cu spreader; 72.67 K and 76.30 K, respectively. Fig. 17 displays the calculated temperature contours in the Cu, composite and PG spreaders cooled by saturation nucleate boiling of FC-72 liquid. The results shown are for a 10  10 mm underlying chip with a non-uniform heat dissipation, Umax = 1.423 (Fig. 13).

The contour lines in the images of Fig. 17 indicates the temperatures in degrees Kelvin. The color1 code is the same for all images. The maximum surface temperature of chip with a Cu spreader (Fig. 17a) is 9 K hotter than with either a composite or a PG

1 For interpretation of color in Figs. 1–18, the reader is referred to the web version of this article.

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thermal resistance of 0.42 °C/W; compared to 48.4 W and only 0.22 °C/W with the composite spreader. The PG spreader removes far less thermal power (24.8 W) and has the highest total thermal resistance (0.46 °C/W). Fig. 18 compares the values of the thermal powers dissipated by the underlying chip and removed by the Cu, PG and the PG–Cu composite spreaders cooled with saturation and 30 K subcooled boiling of FC-72 dielectric liquid. The results show that: (a) Increasing the chip’s heat flux peaking, Umax, or the non-uniformity of the surface heat flux decreases the removed thermal powers by the various spreaders and reduce their footprint areas. (b) The performance of the composite spreader is the best, followed by the Cu and distance third by the PG spreader. (c) The spreaders footprint areas are different, but increase commensurate with the total thermal power removed. (d) The spreaders cooled by subcooled nucleate boiling of FC-72 dielectric liquid remove significantly more thermal powers and typically have 15% smaller footprint areas than the spreaders cooled by saturation boiling of FC-72 liquid. (e) The spreaders remove the largest amount of thermal power dissipated by the underlying 10  10 mm chip when the heat flux at the chip surface is uniform (Umax = 1 and nq = 1). The removed thermal power estimate is as much as 53 W and 100 W with the composite spreader cooled by saturation and 30 K subcooled nucleate boiling of FC-72, respectively (Fig. 18).

Fig. 15. A comparison of the effect of the surface heat flux peaking ratio on the removed thermal powers using PG, Cu and composite spreaders and on the maximum surface temperature of the underlying 10  10 mm chip with nonuniform heat dissipation.

spreader (Fig. 17b and c). In addition, the surface temperature of the Cu spreader is 7.5–9.5 K higher than those of the composite and PG spreaders. The PG spreader has the smallest surface area (2.7 cm2), followed by the composite (7.6 cm2), then the Cu spreader (13.2 cm2) (Figs. 17 and 18). The Cu spreader cooled by saturation boiling of FC-72 dielectric liquid removes 51 W at a total

The results delineated in Fig. 18 clearly show that it is possible to remove up to 100 W of dissipated power by a 10  10 mm chip using a PG–Cu composite spreader cooled by 30 K subcooled nucleate boiling of FC-72 liquid. This is while keeping the total thermal resistance and the chip’s maximum surface temperature lower than with a Cu speeder of the same thickness (2 mm). Because both spreaders are subject to the same cooling constraints (i.e., highest surface hNB is 90% of that at CHF and that at the edge of the spreader corresponds to 1 K higher surface temperature than at boiling incipience), the surface area of the copper spreader is almost twice that of the composite spreader. This is partially because the values of the hNB of FC-72 on PG are much higher than on Cu (Fig. 14b).

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Power Dissipation (W) Power Dissipation (W) Power Dissipation (W)

Fig. 17. Temperature contours in spreaders cooled with saturation boiling of FC-72 when Umax = 1.423 for the 10  10 mm chip with non-uniform heat dissipation.

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Fig. 18. Comparisons of the thermal powers removal by and footprint areas of the composite, PG and Cu spreaders cooled with saturation and 30 K subcooled boiling of FC-72 liquid for 10  10 mm underlying chip with non-uniform heat dissipation.

5. Summary and conclusions This paper presented experimental results of saturation and subcooled nucleate boiling of FC-72 and HFE-7100 dielectric liquid on PG and investigated potential applications of the results to immersion cooling of 10  10 mm chips with non-uniform heat dissipation. Results showed marked enhancements in hMNB and CHF, and lower surface temperatures on PG compared to plane Cu and silicon surfaces. The natural convection data for dielectric liquids on PG, relevant to chip cooling in the standby mode when the surface heat flux is <20 kW/m2, shows 12.4% higher heat removal rates than on plane surfaces. In addition, the values of

hNB and CHF on PG are much higher than those reported by others on plane, micro-porous and structured surfaces and on surfaces with micro-porous coatings. For example, saturation hMNB on PG of 34 kW/m2 K is almost four times that on plane Cu (8.4 kW/ m2 K) and the corresponding surface temperatures are 5–10 K lower. The values of hMNB for saturation and subcooled nucleate boiling of FC-72 on PG (13–33 kW/m2 K) are >2–3 times those on plane Cu (5.3–8.8 kW/m2 K). Increasing the liquid subcooling increases the thermal power removed by nucleate boiling and CHF, but also increases the surface temperature causing hNB to be much lower than for saturation boiling. Saturation nucleate boiling results of FC-72 and HFE-7100 dielectric liquids indicate the

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highest CHF values on PG, followed by those reported by other investigators on micro-porous coatings and then plane surfaces. Results also show the CHF values of FC-72 and HFE-7100 dielectric liquids on various surfaces to increase linearly with increased liquid subcooling, but at different rates. The value of the subcooling coefficient, CCHF,sub, depends on the surface characteristics and the properties of the boiling liquid. Results of saturation and subcooled nucleate boiling of FC-72 and HFE-7100 liquids on PG showed a strong dependence of CHF on surface inclinations. In the upwardfacing orientation (h = 0°), saturation CHF of FC-72 on PG (30 W/ cm2) is 12% higher than reported on micro-porous coatings (26.8 W/cm2) and 57% higher than the average of the reported values on plane Cu (19.1 W/cm2). CHF decreases slowly with increased surface inclination up to 90° then decreases much faster with further increase in inclination to its lowest value at 180°. Beyond 90° inclination, the rate of decrease in CHF on PG with increased inclination is much smaller than reported on surfaces with micro-porous coatings and on plane Cu. In the 180° orientation (downward-facing), the saturation boiling CHF on PG of 16 W/cm2 is 53% of that measured in the upward-facing orientation (0°), while that reported on micro-porous coatings of 4.9 W/ cm2 is only 18% of that reported on the same surface in the upward-facing orientation. The 3-D thermal analysis to size and quantify the performance of the Cu, PG and PG/Cu composite spreaders for cooling 10  10 mm underlying chips with non-uniform heat dissipation provided promising results. All spreaders are of the same thickness (2 mm), but the composite spreader consists of 0.4 mm PG and 1.6 mmthick Cu substrate. In the analysis, the nucleate boiling heat transfer coefficient at the center of the spreader is limited to 90% of that measured at CHF in the pool boiling experiments with saturation and 30 K subcooled FC-72 dielectric liquid. The nucleate boiling heat transfer coefficient at the edge of the spreaders is that corresponding to 1 K surface superheat in the experiments. The composite spreader, which takes advantage of the good heat spreading property of Cu and nucleate boiling enhancements of FC-72 liquid on porous graphite, provides the best performance. It is capable of removing up to 100 W at a reasonable surface temperature, when the dissipated power by the underlying chip is uniform, and the spreader surface is cooled by nucleate boiling of 30 K subcooled FC-72 liquid. Results also show that increasing the nonuniformity of the dissipated thermal power by the chip decreases the total heat removed from the spreader surface and increases the total resistance, the footprint area of the spreader and the maximum surface temperature at the center of the underlying chip. The PG–Cu composite spreaders keeps the chip maximum surface temperature sufficiently low (<63.5 °C), thus ensuring acceptable junction temperature for the underlying chip. Acknowledgments This research has been supported by the University of New Mexico’s Institute for Space and Nuclear Power Studies. The author acknowledges the contributions to the contents of this paper of his former students: Dr. Jack Parker currently with the US Department of Energy; Dr. Hamed Saber currently with the Canadian National Research Council and Dr. Huseyin Bostanci currently with RINI Technology Inc., FL. References [1] El-Genk MS, Bostanci H. Saturation boiling of HFE-7100 from a copper surface, simulating a microelectronic chip. Int J Heat Mass Transfer 2003;46(10):1841–54. [2] El-Genk MS, Bostanci H. Combined effects of subcooling and surface orientation on pool boiling of HFE-7100 from a simulated electronic chip. J Exp Heat Transfer 2003;16(4):281–301.

217

[3] Chang JY, You SM, Haji-Sheikh A. Film boiling incipience at the departure from natural convection on flat, smooth surfaces. J Heat Transfer 1998;120(2):402–8. [4] McNiel A. Pool boiling critical heat flux in a highly wetting liquid. Masters thesis, University of Minnesota, Minneapolis, MN; 1992. [5] Liu ZW, Lin WW, Lee DJ, Peng XF. Pool boiling of FC-72 and HFE-7100. J Heat Transfer 2001;123(2):399–400. [6] Parker JL, E-Genk MS. Saturation and subcooled boiling of HFE-7100 on pinned surfaces at different orientations. J Thermophys Heat Transfer 2009;23(2):381–91. [7] Jiang YY, Wang WC, Wang D, Wang BX. Boiling heat transfer on machined porous surfaces with structural optimization. Int J Heat Mass Transfer 2001;44(2):443–56. [8] Webb RL. Odyssey of the enhanced boiling surface. J Heat Transfer 2004;126(6):1051–9. [9] Rajalu KG, Kumar R, Mohanty B, Varma HK. Enhancement of nucleate pool boiling heat transfer coefficient by reentrant cavity surfaces. J Heat Mass Transfer 2004;41(2):127–32. [10] Vemuri S, Kim KJ. Pool boiling of saturated FC-72 on nano-porous surface. Int Commun Heat Mass Transfer 2005;32(1-2):27–31. [11] El-Genk MS, Parker JL. Nucleate boiling of FC-72 and HFE-7100 on porous graphite at different orientations and liquid subcooling. J Energy Convers Manage 2008;49(4):733–50. [12] Nimkar NK, Bhavnani SH, Jaeger RC. Benchmark heat transfer date for microstructured surfaces for immersion-cooled microelectronics. IEEE Trans Compon Packag Technol 2006;29(1):89–97. [13] Parker JL, El-Genk MS. Enhanced saturation and subcooled boiling of FC-72 dielectric liquid. Int J Heat Mass Transfer 2005;48(18):3736–52. [14] Yu CK, Lu DC, Cheng TC. Pool boiling heat transfer on artificial micro-cavity surfaces in dielectric fluid FC-72. J Micromech Microeng 2006;16(10):2092–9. [15] Arik M, Bar-Cohen A, You SM. Enhancement of pool boiling critical heat flux in dielectric liquids by microporous coatings. Int J Heat Mass Transfer 2007;50(5–6):997–1009. [16] Kim JH, Kashinath MR, Kwark SM, You SM. Optimization of microporous structure in enhancing pool boiling heat transfer of saturated R-123, FC-72 and water. In: Proc ASME-JSME thermal engineering summer heat transfer conference. Paper no. HT2007-32399; 2007. [17] Nakayama W, Daikoku T, Kuwahara H, Nakajima T. Dynamic model of enhanced boiling heat transfer on porous surfaces. Part I: Experimental investigation. J Heat Transfer 1980;102(3):445–50. [18] Anderson TM, Mudawar I. Microelectronic cooling by enhanced pool boiling of a dielectric fluorocarbon liquid. J Heat Transfer 1989;111(3):752–9. [19] Misale M, Guglielmini G, Frogheri M, Bergles AE. FC-72 pool boiling from finned surfaces placed in a narrow channel: preliminary results. J Heat Mass Transfer 1999;34(6):449–52. [20] Rainey KN, You SM. Pool boiling heat transfer from plain and microporous, square pin-finned surfaces in saturated FC-72. J Heat Transfer 2000;122(3):509–16. [21] Jung J-Y, Kwak H-Y. Effect of surface condition on boiling heat transfer from silicon chip with submicron-scale roughness. Int J Heat Mass Transfer 2006;49(23–24):4543–51. [22] Ramaswamy C, Joshi Y, Nakayama W, Johnson WB. Effects of varying geometrical parameters on boiling from microfabricated enhanced structures. J Heat Transfer 2003;125(1):103–9. [23] El-Genk MS, Parker JL. Enhanced boiling of HFE-7100 dielectric liquid on porous graphite. J Energy Convers Manage 2005;46(15–16):2455–81. [24] Wei JJ, Honda H. Effects of fin geometry on boiling heat transfer from silicon chips with micro-pin-fins immersed in FC-72. Int J Heat Mass Transfer 2003;46(21):4059–70. [25] Al-Hajri E, Ohadi M, Dessiatoun SV, Qi J. Thermal performance of microstructured evaporation surfaces – application to cooling of high flux microelectronics. In: Proc ASME international mechanical engineering congress and exposition; 2005. p. 799–805. [26] Ferjancˇicˇ K, Rajšelj D, Golobicˇ I. Enhanced pool boiling CHF in FC-72 from a modulated porous layer coating. J Enhanc Heat Transfer 2006;13(2):175–84. [27] Launay S, Fedorov AG, Joshi Y, Cao A, Ajayan PM. Hybrid micro-nano structured thermal interfaces for pool boiling heat transfer enhancement. Microelectron J 2006;37(11):1158–64. [28] Yu CK, Lu DC. Pool boiling heat transfer on horizontal rectangular fin array in saturated FC-72. Int J Heat Mass Transfer 2007;50(17–18):3624–37. [29] Ujereh S, Fisher T, Mudawar I. Effects of carbon nanotube arrays on nucleate pool boiling. Int J Heat Mass Transfer 2007;50(19–20):4023–38. [30] Ghiu C-D, Joshi YK. Boiling performance on single-layered enhanced structures. J Heat Transfer 2005;127(7):675–83. [31] Watwe AA, Bar-Cohen A, McNiel A. Combined pressure and subcooling effects on pool boiling from a PPGA chip package. J Electron Packag 1997;119(2):95–105. [32] Chang JY, You SM. Boiling heat transfer phenomena from micro-porous and porous surfaces in saturated FC-72. Int J Heat Mass Transfer 1997;40(18):4437–47. [33] El-Genk MS, Saber HH. Composite spreader for cooling computer chip with non-uniform heat dissipation. IEEE Trans Compon Packag Technol 2008;31(1):165–72. [34] Kubo H, Takamatsu H, Honda H. Effects of size and number density of microreentrant cavities on boiling heat transfer from a silicon chip immersed in degassed and gas-dissolved FC-72. J Enhanc Heat Transfer 1999;6(2–4):151–60.

218

M.S. El-Genk / Energy Conversion and Management 53 (2012) 205–218

[35] Rainey KN, You SM. Effects of heater size and orientation on pool boiling heat transfer from microporous coated surfaces. Int J Heat Mass Transfer 2001;44:2589–99. [36] Parker JL, El-Genk MS. Effect of surface orientation on nucleate boiling of FC-72 on porous graphite. J Heat Transfer 2006;128(11):1159–75. [37] El-Genk MS, Parker JL. Nucleate boiling of FC-72 and HFE-7100 on porous graphite at different orientations and liquid subcooling. J Energy Convers Manage 2008;49(4):733–50. [38] El-Genk MS, Saber HH, Parker JL. Efficient spreaders for cooling high performance computer chips. J Appl Therm Eng 2007;27(5–6):1072–88. [39] Chang JY, You SM. Heater orientation effects on pool boiling of microporous enhanced surfaces in saturated FC-72. J Heat Transfer 1996;118:937–43. [40] Rainey KN, You SM, Lee S. Effect of pressure, subcooling, and dissolved gas on pool boiling from microporous surfaces in FC-72. J Heat Transfer 2003;125:75–83. [41] Reed SJ, Mudawar I. Elimination of boiling incipience temperature drop in highly wetting fluids using spherical contact with a flat surface. Int J Heat Mass Transfer 1997;42:2439–54.

[42] Howard AH, Mudawar I. Orientation effects on pool boiling critical heat flux (CHF) and modeling of CHF for near-vertical surfaces. Int J Heat Mass Transfer 1999;42:1665–88. [43] Howard AH. Effects of orientation and downward-facing convex curvature on pool boiling critical heat flux. PhD dissertation, Purdue University, West Lafayette, IN; 1999. [44] Priarone A. Effect of surface orientation on nucleate boiling and critical heat flux of dielectric fluids. Int J Therm Sci 2005;44:822–31. [45] Honda H, Takamastu H, Wei JJ. Enhanced boiling of FC-72 on silicon chips with micro-pin-fins and submicron scale roughness. J Heat Transfer 2002;124:383–90. [46] Arik MA, Bar-Cohen A. Ebullient cooling of integrated circuits by novec fluids. In: Proc Pacific rim intersociety, electronics packaging conference, Kauai, Hawaii; July 18–23, 2001. [47] Kline SJ. The purposes of uncertainty analysis. Trans ASME J Fluids Eng 1985 [June issue].