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Visualization experiments on flat-plate heat pipes with composite mesh-groove wick at different tilt angles
International Journal of Heat and Mass Transfer 123 (2018) 839–847
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International Journal of Heat and Mass Transfer journal homepage: www.elsevier.com/locate/ijhmt
Visualization experiments on flat-plate heat pipes with composite mesh-groove wick at different tilt angles Shwin-Chung Wong ⇑, Wei-Siang Liao 1 Department of Power Mechanical Engineering, National Tsing Hua University, Hsin-Chu 300, Taiwan, ROC
a r t i c l e
i n f o
Article history: Received 4 December 2017 Received in revised form 8 March 2018 Accepted 8 March 2018
Keywords: Heat pipe Evaporator resistance Composite wick Mesh-groove wick Mesh wick Groove wick
a b s t r a c t Visualization experiments are performed on flat-plate heat pipes with a sintered composite copper meshgroove wick at different tilt angles. Deionized water is used as the working fluid. The composite wick consists of a layer of 200-mesh copper screen sintered over parallel semi-circular grooves with a width of 0.18 mm and depth of 0.09 mm. Also investigated are a groove wick with the same groove size and a sintered double-layer 200-mesh (2 200 mesh) wick. The effective length of the heat pipes is 101 mm. The heat pipe performance of the composite mesh-groove wick excels the other two wicks in the maximum heat load (Qmax) under the horizontal orientation and in the anti-gravity ability. Visualization reveals two evaporation stages for the horizontal orientation before full dryout in the heated zone of the composite wick. In the first stage with Q up to about 40 W, the wick at the heated zone is filled with water; in the second stage, partial dryout in the grooves occurs and expands with increasing heat load. Up to about 60 W, the heated zone becomes fully dried. In contrast, the Qmax of the 2 200 mesh wick is 21–25 W, and full dryout prevails at 14 W for the groove wick. When the tilt angle is between 30° and 90°, the Qmax for the composite wick may reach 39–49 W, followed by drastic increase in the evaporator resistance. The 2 200 mesh wick suffers serious dryout at Q = 14 W under a = 45°; the groove wick cannot operate under a 30°. No nucleate boiling is observed in all the tests for water. Ó 2018 Elsevier Ltd. All rights reserved.
1. Introduction Different types of wick are used for heat pipes. The most common wick types are sintered powders, metal meshes, and parallel grooves, each having its own advantages and disadvantages. Among the three, sintered powders have the highest capillarity but lowest permeability, while the groove wicks have the lowest capillarity but highest permeability. To attain a higher maximum heat transfer rate (Qmax, the maximum heat load achieved before the onset of dryout) by optimizing between the capillarity and the permeability, different kinds of composite wicks, such as sintered-powder/groove wicks [1–3] and mesh-groove wicks [4–9], among others [10], have been investigated. Tang et al. [1] verified the higher capillarity of composite sintered-powder/ groove wicks than pure sintered-powder or pure groove wicks. Wong and Chen [2] tested for a composite-wick heat pipe with powders sintered in the grooves at the evaporator section. Both the Qmax and the evaporator resistance were found significantly improved over the heat pipe with a pure groove wick. Li et al. [3] ⇑ Corresponding author. 1
E-mail address: [email protected] (S.-C. Wong). Graduate Student.
https://doi.org/10.1016/j.ijheatmasstransfer.2018.03.031 0017-9310/Ó 2018 Elsevier Ltd. All rights reserved.
examined the performances of heat pipes with composite wicks made by sintering copper powders over parallel grooves. When the grooves were fine enough that the powders were laid outside, the grooves functioned as the liquid passage with low flow resistance. The thermal resistances for fine grooves were lower than those for powder-filled coarse grooves. Considering the potential combinative advantages of high permeability of the grooves and high capillarity from the meshes, heat pipes with a composite mesh-groove wick were first tested by Lefèvre et al. [4]. A layer of fine woven mesh was plated, rather than sintered, against the top walls of parallel rectangular grooves by a coarse mesh. Methanol was used as the working fluid. Unexpectedly, this composite wick resulted in considerably lower Qmax and higher thermal resistance than the groove wick. At a tilt angle of 15°, the heat pipe with this composite wick can work for rather small heat fluxes. The degradation was ascribed to the additional thermal resistance of the mesh layers. The association of the meshes and grooves induced nucleation boiling so that the heat pipe performance turned from capillary-limit to boiling-limit. Hsieh et al. [5] tested flat-plate heat pipes with mesh-covered rectangular grooves with a depth of 0.55 mm and a width of 0.5 mm. One, two, or three layers of 70 lm-thick mesh was plated on the grooves with a supportive structure. The working fluid was acetone. The results for the
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Nomenclature Ac Ae Ah Ap d k Leff N Dpc q Q Qcond Qmax Qps Qt Rc Re Re,min rp t T T
cooled area [cm2] heated area [cm2] cross-sectional area of the heating block [cm2] cross-sectional area of the base plate [cm2] mesh wire diameter [mm] thermal conductivity [W/m °C] effective length of the heat pipe [mm] mesh number [1/mm] capillary pressure [Pa] net heat flux [W/cm2] net heat load (= Qt Qcond) [W] lateral conduction through copper plate [W] maximum heat load before dryout [W] heat input from the power supply [W] total heat load [W] condenser resistance [°C-cm2/W] evaporator resistance [°C-cm2/W] minimum evaporator resistance [°C-cm2/W] pore radius of the wick wick thickness [mm] temperature [°C] thermocouple
horizontal orientation showed that the composite wick with a single mesh layer could improve thermal performance marginally at a small heating area but not at a larger area. With two or three mesh layers, the maximum heat transfer rate decreased and the thermal resistance increased. Under tilt angles, although the composite wicks could maintain heat pipe operation, the thermal performance was seriously degraded. In developing a polymer-based flat-plate heat pipe, Oshman et al. [6] fabricated a copper composite mesh-micropillar wick by electroplating a 200-mesh woven screen on top of micropillars. The 200-lm wide inter-pillar cross-shaped channels were used as the liquid passage with low flow resistance. This polymer-based heat pipe was able to operate at a heat flux of 11.94 W/cm2. Oshman et al. [7] subsequently reduced the width of the inter-pillar channels to 31 lm to form sharp corners and promote capillarity. This modified flat-plate heat pipe with a small footprint (40 mm 40 mm) was demonstrated to operate under adverse acceleration up to 10 g. To characterize the enhanced capillary evaporation with visualization, Dai et al. [8] tested a composite mesh-groove wick vertically placed in a chamber under atmospheric pressure. The composite wick was made by sintering one layer of 145-mesh copper woven screen on top of parallel grooves with a 250 lm 250 lm cross-section. The center of the heating area was 15 mm above the water level. The critical heat flux (CHF) for the composite wick was substantially increased over that for a groove or a mesh wick. The primary reasons for the increased CHF were ascribed to the increased capillarity from the mesh pores and reduced flow resistance of the grooves. Dai et al. [9] further investigated the effects of superhydrophilic treatment on the composite wick using the same test apparatus. The heat transfer coefficient was raised due to enhanced thin film evaporation on the mesh. However, the maximum heat flux was not significantly improved as a result of compromise between the enhanced capillarity and increased viscous drag by the coating materials. At low heat fluxes, oscillating flow owing to the growth and collapse of the boiling bubbles was observed, and the heat transfer coefficient was increased by the induced advection. At larger heat fluxes, the channels became partially dry with vapor cores and the capillary force was mainly provided by the sharp mesh-wire corners. The copper meshes provide larger
x y
direction along the heat pipe direction along the heating block
Greek
a d
e h h0
r x
tilt angle [°] thickness of the immersed part of the wick [m] wick porosity apparent contact angle at the evaporating contact lines [°] static contact angle [°] surface tension [N/m] measurement uncertainty
Subscripts a adiabatic section avg average c condenser cu copper e evaporator eff effective v vapor w wick
surface areas and more nucleation sites for heat transfer. The growth and collapse of bubbles intensified the capillary liquid motion on the meshes and enhanced the thin film evaporation thereon. The works of Oshman et al. [6,7] and Dai et al. [8,9] have verified the advantages of the composite mesh-groove wicks. However, the advantages have not been realized in heat pipes by Lefèvre et al. [4] and Hsieh et al. [5]. The failure was likely because the meshes were plated by supportive structures over the grooves without sintering or electroplating. There could be high contact resistances and possible local detachments to deteriorate the capillarity. It is noted that the visualization studies of Dai et al. [6,7] conducted for water were under the atmospheric condition. The nucleate boiling was observed to be significantly influential to the capillary evaporation process. However, under the low pressures in operating heat pipes, whether boiling occurs and other evaporation characteristics for composite mesh-groove wicks needs be observed by visualization experiments. Heat pipe experiments with visualization have been conducted for sintered-mesh wicks [11–15], sintered-powder wicks [16], groove wicks [17–21], and composite sintered-power/groove wicks [2]. For copper/water heat pipes, nucleate boiling was not observed for thin sintered-mesh wicks [11–15], sintered-powder wicks [16], groove wicks [18], and composite sintered-power/groove wicks [2]. When methanol or acetone was used as the working fluid, weak nucleate boiling occurred near the Qmax which appeared to be still limited by wick capillarity [13]. Hong et al. [22] verified that bubble nucleation is harder to occur in a thin (<1 mm) sintered-powder wick at low pressures. Their theoretical analysis showed that, mainly because of the low vapor density, the wall superheats required for onset of nucleate boiling at low pressures are several times larger than at the atmospheric condition. Therefore, visualization experiments are needed to explore the evaporation characteristics of operating heat pipes with a composite mesh-groove wick which is potentially advantageous in both high permeability and high capillarity. The following questions are worthy to be examined: (1) How is the quantitative thermal performance, namely the maximum heat load and the thermal resistance, of this composite wick compared with
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a mesh-wick and a grooved heat pipe? (2) Could its advantages sufficient to maintain heat pipe operation at adverse tilt angles? (3) Would nucleate boiling occur and influence heat pipe performance at the low-pressure conditions in heat pipes? The present work aims to answer these questions with visualization experiments for a copper/water with a sintered composite mesh-groove wick. For comparison, a heat pipe with the composite wick along with a grooved heat pipe and a double-layer meshed heat pipe will be carefully tested the under horizontal orientation. Then, the mesh-wicked heat pipe and the composite-wicked heat pipe will be tested at various tilt angles. The maximum heat transfer rates, the evaporator resistances, and the visualized evaporation characteristics for these three different wicks will be compared. 2. Experimental methods Fig. 1 shows the overall test setup. The evaporation characteristics of a flat-plate heat pipe were visualized through its glass window using a CCD camera equipped with a microscopic lens (Optem, Zoom 125, resolution up to 600 lp/mm) to capture magnified images, with illumination from a couple of LED lights. Fig. 2 shows the detailed construction of the heat pipe and the thermocouple locations. A wicked base plate of the heat pipe is shown in Fig. 3. The wicked region is 126 mm 50 mm, and the effective length (Leff = 0.5 Le + La + 0.5 Lc, where Le, La, and Lc are the length of the evaporator, adiabatic section, and the condenser, respectively) of the heat pipe is 101 mm. The three different wicks tested in the present work are: a parallel-groove wick, a sintered mesh-groove wick, and a sintered double-layer 200 mesh wick. The crosssectional views of the parallel-groove wick and the mesh-groove
841
Fig. 3. The wicked base plate.
wick are shown in Fig. 4. The grooves were chemically etched, by a commercial etching company, on a 4 mm-thick C1020 oxygenfree copper base plate. The grooves have a roughly semi-circular cross-sectional shape with a width of 0.18 mm and depth of 0.09 mm (Fig. 4a). The sintered composite mesh-groove wick was made by sintering a layer of 200 mesh screen on top of the grooves under a load of 1 kgf/cm2 in a sintering oven with an 800 °C hydrogen/ nitrogen atmosphere for 2 h. Before sintering, the grooved base plate and the mesh screens were carefully cleaned with acetone and water for several times. After sintering, the total depth of the composite sintered wick is 0.18 mm, with a 0.09 mm depth of each portion (Fig. 4b). The average porosity of the composite sintered wick is determined by averaging the groove and the mesh parts. The groove part is 0.50, measured as the cross-sectional area fraction of the grooves. The mesh part is determined using the relation adapted from that for a single-layer 200 mesh wick [10]:
Fig. 1. Experimental setup.
Filling tube
Fig. 2. The flat-plate heat pipe with thermocouple layout.
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(a)
(b)
(c)
0.15mm
Fig. 4. Cross-sectional views of (a) the groove wick, (b) the sintered composite mesh-groove wick, and (c) the 2 200 mesh wick.
e ¼ 1 2:1pNd2 =t; where N is the mesh number, d is the wire diameter, and t is the wick thickness after sintering. With N = 7.874/mm, d = 0.055 mm, and t = 0.09 mm, the porosity of the mesh part is 0.56. Therefore, the porosity of the composite mesh-groove wick is 0.53. The double-layer mesh wick (named as 2 200 mesh wick) was made by sintering two layers of 200 mesh screens on a plain base plate under the same sintering condition as for the composite wick. Its wick porosity e is determined using the following relation for a double-layer mesh wick:
e ¼ 1 1:05pNd2 =t: With t = 0.15 mm after sintering, as shown in Fig. 4c, e = 0.48. The static contact angles for various surface conditions were measured right before the assembly of the heat pipe using a contact angle meter (Mactech Co.). The measurements were made for sessile drops of 5 ll at three locations on the flat edge region of the copper base plate. The static contact angles were h0 = 13 ± 1°, in agreement with our previous measurements [2,12–15,18]. The present contact angle values are smaller than most of the literature values [23–25], which scatter between 7° and 94° and are sensitive to surface roughness, cleansing process and oxidation condition. As our copper surfaces were carefully cleaned and reduced in a hightemperature hydrogen/nitrogen atmosphere, low values have been obtained consistently. To keep its h0 within the range of 13 ± 1°, a
reduction process was necessary for each test run. The reduction process was similar to the sintering process except that no pressure was exerted on the base plate. To avoid degradation of the contact angle, evacuation of the assembled heat pipe started no more than 5 min after the base plate was taken out of the sintering oven. Also, after the tests the base plate could maintain good wettability. The inner space of the heat pipe was 126 mm 50 mm 5 mm. With the interfaces between different pieces sealed with orings, the whole structure, including the wicked copper base plate, the upper glass window, and stainless-steel frames, was tightened with bolts. A uniform heat load Qt, through a copper block, was supplied by a cartridge heater embedded in another copper block below the upper one. The cartridge heater was powered by a DC power supply. The heating area was 10 mm 50 mm; the watercooled area was 40 mm 50 mm, provided by a cold plate with cooling water of 20 ± 0.5 °C. The interfaces between the heating block/cold plate and the base plate were laid with a layer of thermal grease (Dow Corning TC5021, thermal conductivity of 3.3 W/ mK). The heating block and the cold plate were insulated within a Bakelite box, on which the flat-plate heat pipe was positioned. The heat pipe and the Bakelite box were altogether wrapped with ceramic wool except for short periods of video recording. Since most of the tests were conducted with the adiabatic-section temperatures near the room temperature, condensation on the glass window was not serious in our experiments as the heat loss through the window was limited. When the window was indeed blocked by a water layer, a hair blower was used to evaporate the water on the window. Afterwards, thermal equilibrium was re-attained before further measurements. To reduce the longitudinal conduction through the 4 mm-thick copper base plate, trenches were made across the plate and around the evaporator and the condenser to leave a local plate thickness of 0.8 mm (Figs. 2 and 3). The fluid charge was somewhat more than the saturate amount to fill the wick voids. The excess charge was needed to compensate for the liquid trapped at the corners, in the gaps, and on the walls of the heat pipe. Nineteen K-type SS-sheathed probe thermocouples (Omega, Inc.) were used to measure the temperatures at selected positions. Among them, these measuring the internal vapor temperatures were inserted via holes drilled through the middle stainless-steel frame. To prevent air leakage, vacuum fittings were installed to accommodate these /1.6-mm thermocouples. All the other probe thermocouples were /1.0 mm. Connected to a data logger (Fluke, Hydra Series II), they yielded temperature readings with a resolution of 0.1 °C. The readings of these thermocouples have been checked to agree well with each other under the room temperature and the boiling water. As shown in Fig. 2, thermocouples T1, T2 and T3, each 5 mm apart, were embedded in the upper heater block to measure Qt. Calculations used T1 and T3 (T2 as a reference for linearity checking) and the relation (Dy = 10 mm)
Q t ¼ kcu Ah ðT 1 T 3 Þ=Dy;
ð1Þ
where kcu is the thermal conductivity of copper, determined from the standard physical property data, and Ah is the cross-sectional area of the heating block. Thermocouples T4–T6 measured temperature T4–T6 of the copper plate in the heated area, with T5 at the center. T7 was positioned at 1 mm above the center of the evaporator to reflect the vapor temperature T7 leaving the evaporator. They were used to determine the evaporator resistance. The lateral conduction through the copper plate, Qcond, was estimated using a pair of thermocouples embedded in the copper base plate (T9 and T10) with
Q cond ¼ kcu Ap ðT 9 T 10 Þ=Dx;
ð2Þ
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where Ap is the cross-sectional area of the base plate and Dx is the distance between thermocouples T9 and T10. In general, these calculated ratios of Qcond/Qt were about 3–5% prior to the onset of partial dryout in the composite-wick evaporator. The net heat load Q was determined by
Q ¼ Q t Q cond :
ð3Þ
Thermocouples T11, T12, and T13, 10 mm apart, measured the base plate temperatures T11, T12, T13 below the condensation zone. All the thermocouples for base-plate temperature measurement (T4–T6 and T9–T13) were positioned along the central axis of the heat pipe. At 1 mm above the condenser wick were T14, T15, and T16, which measured the vapor temperatures needed to calculate the condenser resistance. Furthermore, 4 mm above the condenser wick were placed T17, T18, and T19 to monitor the extent of noncondensable gas (NCG). The NCG might have entered along with the charge water or leak in through the o-ring sealing during the test. To reliably measure the heat pipe performance, NCG was carefully minimized in the charging process and monitored throughout the tests. De-ionized water was used as the working fluid. Before each test, the working fluid was degassed by boiling in a long-neck flask, whose outlet is connected with a Graham condenser, for a sufficient duration. Right after the heat pipe was evacuated down to a pressure of 8 103 Torr, a selected volume of degassed liquid was filled into the heat pipe via a filling tube welded on the stainless-steel frame (Fig. 2). Afterwards, the filling tube was shut by a valve. The heat pipe was then set up for testing. The heating from the DC power supply began from a low power and was incremented stepwise. Except for a special test run, the duration of each test was within 4–6 h. Data were taken under a thermally stable condition for each heat load, with all temperature variations less than 0.2 °C within the last 10 min. Since the dryout behaviors are quite different for the three types of wick, the tests proceeded until the base temperatures at the evaporator increase drastically, reflecting partial or even full dryout in the evaporator. The Qmax was then determined as the maximum heat load before the drastic increase in Re. Since the small amount of water trapped at the corners, gaps and walls varied in different tests, the results of Qmax, Re, and especially the condenser resistance Rc for the same test conditions may vary. The evaporator resistance Re and the condenser resistance Rc were determined respectively as
Re ¼ ðT e;av g T 7 Þ=qe Rcu ½ C-cm2 =W
ð4Þ
Rc ¼ ðT v ;av g T c;av g Þ=qc Rcu ½ C-cm2 =W
ð5Þ
where Te,avg = (T4+2T5+T6)/4, Tv,avg = (T14+T15+T16)/3, Tc,avg = (T11 + T12 + T13)/3, based on the portion of coverage of each thermocouple in the measured region; Rcu (= 0.0875 °C-cm2/W) accounts for the thermal resistance associated with the distance of 3.5 mm between the wick bottom and the thermocouples embedded in the grooves made on the lower surface of the copper plate. Since the heat pipe was well insulated during the tests, the heat loss to the environment was neglected. Hence the heat Q via evaporation was considered equal to the heat via condensation. In Eqs. (4) and (5), the net heat fluxes qe = Q/Ae, and qc = Q/Ac, with the heated area Ae = 5 cm2 and the cooled area Ac = 20 cm2. The experimental uncertainties were calculated using the propagation-of-error method. As the power supply input was calculated by Qps = I V, the relative uncertainty of Qps can be determined as
x Q ps Q ps
" #1=2 xðIÞ 2 xðVÞ 2 ¼ þ : I V
ð6Þ
With x(I) = ±0.005 Amp and x(V) = ±0.05 Volt, the uncertainty of Qps was less than ±1%. The relative uncertainty of Qt was determined according to Eq. (1) as
xðQ t Þ Qt
" ¼
xðkcu Þ kcu
2
þ
xðAh Þ Ah
2
þ
2
xðDTÞ DT
2 #1=2
þ
xðDyÞ
:
Dy
ð7Þ The relative uncertainty for kcu was ±0.5% and those for Ah and Dy were negligible. In temperature measurements, the thermocouple readings were rather stable. The average of these readings lay within an uncertainty of ±0.05 °C for a single thermocouple. Thus, the uncertainty in DT associated with two thermocouples was ±0.07 °C. For the composite wick, the operation range of Q was 13.4–66 W. For 30 W < Q < 66 W, the values of (T1 T3) used in Eq. (1) to calculate Qt ranged between 1.6 and 4.0 °C. Hence, the relative uncertainty in Qt was between ±2 and 5%, with a larger uncertainty at a lower Qt. For 17 W < Q < 30 W, (T1 T3) ranged between 0.9 and 1.6 °C. Hence, the relative uncertainty in Qt was between ±5 and 8%. The maximum uncertainty in Qt was ±10% at the lowest Q of 13.4 W. The relative uncertainty in Q (= Qt Qcond) was near that of Qt since Qcond was only a small fraction of Qt. For the 2 200 mesh wick, the Q range was 12–26 W with (T1 T3) between 0.7 and 1.3 °C. Hence, the uncertainty in Q was between ±5 and 10%. For the groove wick, Q ranged between 9.4 and 14.1 W with (T1 T3) between 0.6 and 1.1 °C, and the uncertainty in Q was between ±6 and 12%. The uncertainties in Re and Rc were calculated using the propagation of error method based on Eqs. (4) and (5), respectively. For the composite wick, the relative uncertainties in Re were ±17% or 26% for Q = 13.4 W (two cases); ±9– 14% for 17 W < Q < 30 W; ±2–9% for 30 W < Q < 66 W, and ±4– 14% in Rc with a larger uncertainty at a lower Q. For the 2 200 mesh wick, the uncertainties were ±7–14% in Re and ±7–14% in Rc. For the groove wick, the uncertainties were ±7–15% (but ± 37% at 9.4 W) in Re and ±9–14% in Rc. The measurement uncertainties for heat pipes with different wicks are summarized in Table 1. Another way to roughly estimate the uncertainty in Qt is by comparing the power supply input and the heat-conduction calculation by Eq. (1). In general, their differences are within 15%. 3. Results and discussion 3.1. The composite mesh-groove wick Fig. 5 illustrates the longitudinal temperature distributions in the vapor space and the copper base plate for the horizontal heat pipe with the composite mesh-groove wick at four representative heat loads. The liquid charge v in this test run is 0.9 ml (the saturate charge to fill the wick is 0.6 ml). In this figure, the vapor temperatures above the condenser are represented by T17–T19
Table 1 Summary of measurement uncertainties for different wicks. Composite mesh-groove wick
2 200 mesh wick
Groove wick
Qps
<±1%
Qt or Q
±10%, Q = 13.4 W ±5–8%, 17 W < Q < 30 W ±2–5%, 30 W < Q < 66 W
±5–10%
±6–12%
Re
±17% or ±26%, Q = 13.4 W ±9–14%, 17 W < Q < 30 W ±2–9%, 30 W < Q < 66 W
±7–14%
±7–15% (but ±37% at 9.4 W)
Rc
±4–14%
±7–14%
±9–14%
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φ1.6 mm TC
Fig. 5. Longitudinal temperature distributions in the vapor and the base plate at different heat loads for the horizontal heat pipe with a sintered composite meshgroove wick, v = 0.9 ml.
partial dryout
heated zone
Fig. 6. Image of the evaporator with partial dryout, v = 0.9 ml, a = 90°, Q = 39.4 W.
2
2 The reasons for the differences between the upper and lower thermocouple readings have been discussed in [15]. 3 As Q is increased to 56.5 W or 63.4 W, the vapor temperatures (T17–T19) become lower than that at the adiabatic section (T8), reflecting mild NCG effect. It is noted that the total test time of this specific test run (Case 1 with v = 0.9 ml shown in Fig. 7), with data taken at a large number of heat loads, was about 8 h. Some amount of NCG might have leaked in during the long test time. All other tests were within 4–6 h, and the NCG were controlled within a limited extent.
Re, case1, v = 0.9 ml Re, case2, v = 0.95 ml
composite mesh-groove wick
Rc, case1 , v = 0.9 ml
1.5
R (K∙cm²/W)
measured by the upper thermocouples T17–T19. Among the six thermocouples (T14–T19) above the condenser (cf. Fig. 2), the lower three (T14–T16) 1 mm above the condenser wick were used for Rc calculation (Eq. (5)), while the upper three (T17–T19) 4 mm above the wick were used to monitor the NCG effect. The present results are in similar trends as those obtained in our previous measurements [14,15].2 At Q = 19.5 W and 42.2 W, the vapor temperatures are rather uniform even near the end of the condenser region, showing negligible NCG accumulation near the condenser end during the early period of the test.3 This feature also evidences the insignificant longitudinal pressure drop and heat loss (either through the upper window or to the base plate) associated with the vapor flow across the upper space of the chamber. At Q = 56.5 W, partial dryout has occurred in the heated zone so that the plate temperatures rise drastically. In addition, the left-end thermocouple reading is the highest, with T4 > T5 > T6 because partial dryout initiates from the end of the evaporator region. The image in Fig. 6 reflects the uneven distribution of the partial-dryout region over the heated zone for the composite mesh-groove wick. When partial dryout occurs in certain grooves, the hot dryout region would prompt the onset of partial dryout in neighboring grooves. Besides, the dryout lengths in individual grooves are different. With the heat load increased, the dryout regions expand until the whole heated zone dries out, yielding high plate temperatures. The condition at Q = 63.4 W presented in Fig. 5 corresponds to the situation when the dryout region almost covers the whole heated zone. At Q = 56.5 W (partial dryout) and 63.4 W (near full dryout), the vapor temperatures measured above the heated zone (T7) are higher than those at the adiabatic section (T8). This is because thermocouple T7 is influenced by the upstream vapor above the dryout region. After sufficient mixing, the vapor temperature settles to the equilibrium temperature of T8. Another point to note is that the slight drop of the plate temperature at the condenser end (T13) is because the cooling water enters the cold plate from the end side. Fig. 7 presents the evaporator and condenser resistances for two horizontal test runs for the composite-wick heat pipe. Fig. 8 shows the magnified images of the evaporator at different heat loads,
Rc, case2, v = 0.95 ml
1
0.5
0 0
10
20
30
Q(W)
40
50
60
70
Fig. 7. Evaporator and condenser resistances versus Q for two test runs for the composite wick.
with the images taken from Case 2 (v = 0.95 ml) in Fig. 7. For both cases, the evaporator resistances (Re) remain around a fixed value up to Q 43 W. In this stage, the evaporator remains fully wetted, as shown in Fig. 8a, thanks to the capillary force provided by the small mesh pores. When Q > 43 W, partial dryout initiates from the upstream end of the heated zone. The dryout region expands with increasing Q, leading to gradually increased Re, as shown in Case 2. Fig. 8b shows two neighboring grooves with one dried and another wetted. During this stage, the values of Re for Case 1 remains near a constant lower than for Case 2. But for a third run not shown here, the variation trend of Re is similar to that of Case 2 with gradually rising Re during the partial-dryout stage. The reason why the Res of Case 1 are distinctly lower than those of the other two cases is unclear. When Q is further increased beyond 60 W or 63 W, another jump in Re appears. Fig. 8c shows the image taken around the downstream edge of the heated zone at Q = 66 W, when the dryout region almost covers the whole heated zone. At this heat load, Re is stably maintained at a high value. Note that at this test condition, Qcond/Qt increases to 12% as a result of the high evaporator temperature. In comparison, Qcond/Qt remains less than 5% in the fully-wetted stage and between 5 and 9% in the partial-dryout stage. As far as the condenser resistances Rc are concerned, the present values are in similar trend as our previous experimental results [14,15] that Rcs are several times larger than Res. Note that the Rc data are sensitive to the amount of liquid trapped in the internal gaps or on the walls of the heat pipe. Since the trapped amount varied in different tests and different times in the same test, the Rc data can only be regarded as qualitative reference.
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Fig. 8. Magnified images of the evaporator, (a) fully wetted at Q = 29.8 W, (b) partial dryout at Q = 47.5 W, (c) near full dryout at the downstream edge of the heated zone, Q = 66.0 W.
Fig. 9 compares the evaporator resistances for a composite-wick heat pipe under various tilt angles. All the cases, horizontal or tilted, exhibit nearly unvaried Res over the first stage up to 39– 49 W. During the first stage, the heated zone is largely fully wetted. This is because when the grooves are covered with a mesh layer, they function as capillary tubes with the mesh pores providing intense capillary force. Thus, dryout within the grooves is postponed and good anti-gravity ability can be gained. Even at a high tilt angle of 90°, the low-Re stage can be maintained up to 39.4 W. Beyond the first stage, different Re–Q trends are observed for the two a = 30° cases. In Case 1, Re rises gradually as Q > 40 W, while in Case 2, Re remains low up to Q = 49 W, followed by a drastic jump. The gradual Re rise in the former case is similar to Case 2 of a = 0°, during which partial dryout expands gradually over the
1
Re ( C∙cm /W)
φ1.6 mm TC
0° case 1 , v = 0.9 ml 0° case 2 , v = 0.95 ml 30° case 1 , v = 0.9 ml 30 ° case 2, v = 0.9 ml 45 ° , v = 0.95 ml 90 °, v = 0.9 ml
0.8
0.6
heated zone. Drastic jump at the end of the first stage occurs for a 45°. The partial-dryout stage, which sustains substantially at a = 0°, stays within a narrow range of heat load. The narrow transitional partial-dryout stage is illustrated using Figs. 6 and 10 for a = 90°. In Fig. 6, at Q = 39.4 W when partial dryout has first appeared, about 1/3 of the heated zone has dried out. But when Q is increased to 43.4 W, the dryout region expands to cover almost the whole heated zone, as shown in Fig. 10. Note that the elapse time for the expansion to stabilize was as long as 40 min. For high tilt angles, once partial dryout appears, it could develop to full dryout of the heated zone after a mild increase of heat load, leading to a jump in Re. In contrast, the dryout expansion process for a = 0° was gradual and the elapse times for stabilization were about half as long as for a = 90°.
0.4
dryout
heated zone
wetted
0.2
0
0
10
20
30
40
50
60
70
Q (W) Fig. 9. Evaporator resistances versus Q for the composite wick under various tilt angles.
Fig. 10. Image of the evaporator at a = 90° with the heated zone nearly fully dried at Q = 43.4 W.
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3.2. The groove wick Fig. 11 illustrates the variations of Re and Rc versus Q for a horizontal grooved heat pipe with v = 0.6 ml (the saturate charge to fill the groove volume is 0.3 ml). At Q = 9.4 W, the groove wick is fully wetted, as shown in Fig. 12. When Q is increased to 11.6 W, partial dryout can be observed. At Q = 13.2 W, the dryout region has extended over the middle line (marked by the thermocouple) of the heated zone, as shown in Fig. 13. When Q = 14.1 W, full dryout has prevailed the heated zone. The zigzag shape of the steepsloped liquid fronts has been observed by Wong and Chen [18].
1.6 groove wick, v = 0.6 ml
Re
1.4
The back-and-forth oscillation of the liquid fronts observed in [18] was not obvious in the present tests. For the case shown in Fig. 11, the liquid fronts were stagnant at a fixed Q. In some other tests, very mild oscillations were observed. The liquid frontline receded with increasing Q, leading to higher temperatures in the heated zone and hence higher Re. Detailed description of the receding process of the liquid fronts in connection with the continuous temperature increase in the heated zone has been provided in [18]. No boiling was observed, the same as in Wong and Chen for water [18]. For grooved heat pipes, Re rises continuously with increasing Q as a result of continuous liquid-front recession. This has been indicated by the visualization experiments of Wong and Chen [18] and by Lin and Wong [26] for commercial grooved heat pipes. The groove wick cannot operate under tilted condition with a 30°.
Rc
R ( C∙cm²/W)
1.2
3.3. The 2 200 mesh wick
1 0.8 full dryout 0.6 partial dryout 0.4 0.2 0
fully wetted 0
5
10
15
Q (W)
20
Fig. 11. Evaporator resistances versus Q for the horizontal heat pipe with the groove wick.
heated region
Fig. 14 shows the Re and Rc results versus Q for a horizontal heat pipe with the 2 200 mesh wick. The Re data of the four runs are quite consistent. The Qmaxs range around 21–25 W, followed by the onset of dryout and marked increase in Re. The minimum evaporator resistance Re,min occurs around Qmax, as a result of thinnest liquid layer thickness in the evaporator. Detailed description of the dryout process for multi-layer mesh wicks in connection with visualization images and temperature measurements at the evaporator is available in Liou et al. [11] and Wong et al. [13]. Similar as in Liou et al. [11], no boiling was observed here for water. The values of Rc at v = 1.3 ml (the saturate charge is 0.5 ml) are higher than those at v = 1.1 ml due to the thicker liquid layer at the condenser. The condensation mechanisms for a mesh wick with different surface wettabilities, rendered by different working fluids or surface treatments, have been carefully discussed by Wong et al. [14,15]. Tests under a = 30° and 60° were attempted, but the heated zone dried out at Q = 14 W, yielding high Re values as large as 0.66 °C-cm2/W and 1.0 °C-cm2/W, respectively. Under a tilt angle, the wick capillary was insufficient to sustain wetted evaporator against liquid flow resistance and additional gravity at this heat load. 3.4. Comparison between different wicks
thermocouple
Fig. 12. Image of the fully wetted groove evaporator, Q = 9.4 W.
centerline of heated region
liquid front
According to the above heat pipe test results, the sintered composite mesh-groove wick is remarkably superior to the others, due to the combinative advantages of high capillarity from the mesh screen and high permeability of the parallel grooves. For horizontal operation of the composite wick, Qmax may be defined as the
1.8 Re, v = 1.1 ml, case 1
1.6
1.4
R ( C∙cm²/W)
2 200 mesh wick
Rc, v = 1.1 ml, case 1 Re, v = 1.1 ml, case 2
1.2
Rc, v = 1.1 ml, case 2
1
Re, v = 1.3 ml, case 1
Rc, v = 1.3 ml, case 1
0.8
Re, v = 1.3 ml, case 2
0.6
Rc, v = 1.3 ml, case 2
0.4 0.2
1.6 mm wetted
dryout
thermocouple
Fig. 13. Image of the groove evaporator with partial dryout, Q = 13.2 W.
0
0
5
10
15
Q (W)
20
25
30
Fig. 14. Evaporator and condenser resistances versus Q for the horizontal heat pipe with the 2 200 sintered mesh wick.
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maximum Q prior to the full dryout, since Res remain reasonably low. Then, the Qmax of about 60 W is much higher than 21–25 W for the 2 200 mesh wick and about 10 W for the groove wick. Even under a tilt angle of 30–90°, the Qmax may reach 39–49 W. In contrast, the 2 200 mesh wick cannot operate at Q = 14 W under a = 30°. The negative evaluation on the composite mesh-groove wick of Lefèvre et al. [4] was likely because the meshes were plated over the grooves only by supportive structures. There would be high contact resistances and possible local detachments to deteriorate the capillarity. However, in the present study and [6–9], in which the mesh screen was sintered or electroplated onto the grooves, positive results have been obtained. Therefore, secure bonding between the mesh screens and the grooves, either by sintering [8,9] or electroplating [6,7], is essential for the composite meshgroove wicks to manifest their superiority as the heat pipe wick. 4. Conclusions Visualization experiments have been performed on flat-plate heat pipes with a sintered composite copper mesh-groove wick at different tilt angles. For comparison, a groove wick with the same groove size and a sintered 2 200 mesh wick also have been investigated. The effective length of the heat pipes is 101 mm. Deionized water is used as the working fluid. The following conclusions are reached: 1. The composite wick exhibits two evaporation stages under the horizontal orientation before full dryout in the heated zone. In the first stage with Q up to about 40 W, the wick is filled with water; in the second stage, partial dryout in the grooves occurs and expands with increasing heat load. Up to about 60 W, the heated zone becomes fully dried. In contrast, the Qmax of the 2 200 mesh wick is 21–25 W, and full dryout prevails at 14 W for the groove wick. 2. When the tilt angle is between 30° and 90°, the Qmax for the composite wick may reach 39–49 W, followed by drastic increase in the evaporator resistance. The 2 200 mesh wick suffers serious dryout at Q = 14 W under a = 30°; the groove wick cannot operate under a 30°. 3. The sintered composite mesh-groove wick is superior to the multi-layer mesh wick and the groove wick because of its combinative advantages of high capillarity from the mesh screen and high permeability of the parallel grooves. 4. No nucleate boiling is observed in all the present tests for the different wicks with water as the working fluid. The heat pipe performance is capillary-limited. 5. Secure bonding between the mesh screens and the grooves, either by sintering or electroplating, is essential for the composite mesh-groove wicks to manifest their superiority as the heat pipe wick. The present performance tests have verified the superiority of the sintered composite mesh-groove wick. Our future work will measure the permeabilities and the effective pore sizes of these different wicks. Working fluid other than water will also be tested. Until then, the superiority of the sintered composite mesh-groove wick as a result of its combinative advantages of both high capillarity and high permeability can be quantitatively proved. Disclosure statement All authors of this paper have no actual or potential conflict of interest including any financial, personal or other relationships with other people or organizations.
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Acknowledgments This work was funded by Ministry of Science and Technology of Taiwan, R.O.C. under the Project MOST105-2221-E-007-063. The cross-sectional pictures of the groove wick and the mesh-groove wick were taken by Mr. Min-Chieh Liu. References [1] Y. Tang, D. Deng, L. Lu, M. Pan, Q. Wang, Experimental investigation on capillary force of composite wick structure by IR thermal imaging camera, Exp. Therm. Fluid Sci. 34 (2010) 190–196. [2] S.-C. Wong, C.-W. Chen, Visualization experiments for groove-wicked flat-plate heat pipes with various working fluids and powder-groove evaporator, Int. J. Heat Mass Transfer 66 (2013) 396–403. [3] Y. Li, H.F. He, Z.X. Zeng, Evaporation and condensation heat transfer in a heat pipe with a sintered-grooved composite wick, Appl. Therm. Eng. 50 (2013) 342–351. [4] F. Lefèvre, J.B. Conrardy, M. Raynaud, J. Bonjour, Experimental investigations of flat plate heat pipes with screen meshes or grooves covered with screen meshes as capillary structure, Appl. Therm. Eng. 37 (2012) 95–102. [5] J.-C. Hsieh, H.-J. Huang, S.-C. Shen, Experimental Study of microrectangular groove structure covered with multi mesh layers on performance of flat heat pipe for LED lighting module, Microelectron. Reliab. 52 (2012) 1071–1079. [6] C. Oshman, B. Shi, C. Li, R.G. Yang, Y.C. Lee, V.M. Bright, The development of polymer-based flat heat pipes, J. Microelectromech. Syst. 20 (2011) 410–417. [7] C. Oshman, Q. Li, L.A. Liew, R. Yang, Y.C. Lee, V.M. Bright, D.J. Sharar, N.R. Jankowski, B.C. Morgan, Thermal performance of a flat polymer heat pipe heat spreader under high acceleration, J. Micromech. Microeng. 22 (2012) 045018 (12 pp). [8] X. Dai, F. Yang, R. Yang, Y.C. Lee, C. Li, Micromembrane-enhanced capillary evaporation, Int. J. Heat Mass Transfer 64 (2013) 1101–1108. [9] X. Dai, M. Famouri, A.I. Abdulagatov, R. Yang, Y.C. Lee, S.M. George, C. Li, Capillary evaporation on micromembrane-enhanced microchannel wicks with atomic layer deposited silica, Appl. Phys. Lett. 103 (2013) 151602. [10] A. Faghri, Heat Pipe Science and Technology, second ed., Global Digital Press, 2016. [11] J.-H. Liou, C.-W. Chang, C. Chao, S.-C. Wong, Visualization and thermal resistance measurement for the sintered mesh-wick evaporator in operating flat-plate heat pipes, Int. J. Heat Mass Transfer 53 (2010) 1498–1506. [12] S.-C. Wong, Y.-C. Lin, Effect of copper surface wettability on the evaporation performance: Tests in a flat-plate heat pipe with visualization, Int. J. Heat Mass Transfer 54 (2011) 3921–3926. [13] S.-C. Wong, Y.-C. Lin, J.-H. Liou, Visualization and evaporation resistance measurement in heat pipes charged with water, methanol or acetone, Int. J. Thermal Sci. 52 (2012) 154–160. [14] S.-C. Wong, H.-H. Tseng, S.-H. Chen, Visualization experiments on the condensation process in heat pipe wicks, Int. J. Heat Mass Transfer 68 (2014) 625–632. [15] S.-C. Wong, H.-S. Cheng, C.-W. Tu, Visualization experiments on the performance of mesh-wick heat pipes with differing wick wettability, Int. J. Heat Mass Transfer 114 (2017) 1045–1053. [16] S.-C. Wong, J.-H. Liou, C.-W. Chang, Evaporation resistance measurement with visualization for sintered copper-powder evaporator in operating flat-plate heat pipes, Int. J. Heat Mass Transfer 53 (2010) 3792–3798. [17] S. Lips, F. Lefèvre, J. Bonjour, Nucleate boiling in a flat grooved heat pipe, Int. J. Therm. Sci. 48 (2009) 1273–1278. [18] S.-C. Wong, C.-W. Chen, Visualization and evaporator resistance measurement for a groove-wicked flat-plate heat pipe, Int. J. Heat Mass Transfer 55 (2012) 2229–2234. [19] M.J. Stubblebine, I. Catton, Passivation and performance of inorganic aqueous solutions in a grooved aluminum flat heat pipe, J. Heat Transfer 137 (2015) 052901 (8 pp). [20] R. Savino, D. De Cristofaro, A. Cecere, Flow visualization and analysis of selfrewetting fluids in a model heat pipe, Int. J. Heat Mass Transfer 115 (2017) 581–591. [21] P. Xu, Q. Li, Visualization study on the enhancement of heat transfer for the groove flat-plate heat pipe with nanoflower coated CuO layer, Appl. Phys. Lett. 111 (2017) 141609 (5 pp). [22] F.J. Hong, P. Cheng, H.Y. Wu, Z. Sun, Evaporation/boiling heat transfer on capillary feed copper particle sintered porous wick at reduced pressure, Int. J. Heat Mass Transfer 63 (2013) 389–400. [23] M.G. Semena, A.G. Kostornov, A.N. Gershuni, V.K. Zaripov, Contact angles of wicks for low-temperature heat pipes, J. Eng. Phys. 28 (1975) 147–150. [24] M. Yekta-Fard, A.B. Ponter, Surface treatment and its influence on contact angle of water drops residing on Teflon and copper, J. Adhesion 18 (1985) 197– 206. [25] G.S. Hwang, Y. Nam, E. Fleming, P. Dussinger, Y.S. Ju, M. Kaviany, Multi-artery heat pipe spreader: Experiment, Int. J. Heat Mass Transfer 53 (2010) 2662– 2669. [26] K.-T. Lin, S.-C. Wong, Performance degradation of flattened heat pipes, Appl. Therm. Eng. 50 (2013) 46–54.
Contents lists available at ScienceDirect
International Journal of Heat and Mass Transfer journal homepage: www.elsevier.com/locate/ijhmt
Visualization experiments on flat-plate heat pipes with composite mesh-groove wick at different tilt angles Shwin-Chung Wong ⇑, Wei-Siang Liao 1 Department of Power Mechanical Engineering, National Tsing Hua University, Hsin-Chu 300, Taiwan, ROC
a r t i c l e
i n f o
Article history: Received 4 December 2017 Received in revised form 8 March 2018 Accepted 8 March 2018
Keywords: Heat pipe Evaporator resistance Composite wick Mesh-groove wick Mesh wick Groove wick
a b s t r a c t Visualization experiments are performed on flat-plate heat pipes with a sintered composite copper meshgroove wick at different tilt angles. Deionized water is used as the working fluid. The composite wick consists of a layer of 200-mesh copper screen sintered over parallel semi-circular grooves with a width of 0.18 mm and depth of 0.09 mm. Also investigated are a groove wick with the same groove size and a sintered double-layer 200-mesh (2 200 mesh) wick. The effective length of the heat pipes is 101 mm. The heat pipe performance of the composite mesh-groove wick excels the other two wicks in the maximum heat load (Qmax) under the horizontal orientation and in the anti-gravity ability. Visualization reveals two evaporation stages for the horizontal orientation before full dryout in the heated zone of the composite wick. In the first stage with Q up to about 40 W, the wick at the heated zone is filled with water; in the second stage, partial dryout in the grooves occurs and expands with increasing heat load. Up to about 60 W, the heated zone becomes fully dried. In contrast, the Qmax of the 2 200 mesh wick is 21–25 W, and full dryout prevails at 14 W for the groove wick. When the tilt angle is between 30° and 90°, the Qmax for the composite wick may reach 39–49 W, followed by drastic increase in the evaporator resistance. The 2 200 mesh wick suffers serious dryout at Q = 14 W under a = 45°; the groove wick cannot operate under a 30°. No nucleate boiling is observed in all the tests for water. Ó 2018 Elsevier Ltd. All rights reserved.
1. Introduction Different types of wick are used for heat pipes. The most common wick types are sintered powders, metal meshes, and parallel grooves, each having its own advantages and disadvantages. Among the three, sintered powders have the highest capillarity but lowest permeability, while the groove wicks have the lowest capillarity but highest permeability. To attain a higher maximum heat transfer rate (Qmax, the maximum heat load achieved before the onset of dryout) by optimizing between the capillarity and the permeability, different kinds of composite wicks, such as sintered-powder/groove wicks [1–3] and mesh-groove wicks [4–9], among others [10], have been investigated. Tang et al. [1] verified the higher capillarity of composite sintered-powder/ groove wicks than pure sintered-powder or pure groove wicks. Wong and Chen [2] tested for a composite-wick heat pipe with powders sintered in the grooves at the evaporator section. Both the Qmax and the evaporator resistance were found significantly improved over the heat pipe with a pure groove wick. Li et al. [3] ⇑ Corresponding author. 1
E-mail address: [email protected] (S.-C. Wong). Graduate Student.
https://doi.org/10.1016/j.ijheatmasstransfer.2018.03.031 0017-9310/Ó 2018 Elsevier Ltd. All rights reserved.
examined the performances of heat pipes with composite wicks made by sintering copper powders over parallel grooves. When the grooves were fine enough that the powders were laid outside, the grooves functioned as the liquid passage with low flow resistance. The thermal resistances for fine grooves were lower than those for powder-filled coarse grooves. Considering the potential combinative advantages of high permeability of the grooves and high capillarity from the meshes, heat pipes with a composite mesh-groove wick were first tested by Lefèvre et al. [4]. A layer of fine woven mesh was plated, rather than sintered, against the top walls of parallel rectangular grooves by a coarse mesh. Methanol was used as the working fluid. Unexpectedly, this composite wick resulted in considerably lower Qmax and higher thermal resistance than the groove wick. At a tilt angle of 15°, the heat pipe with this composite wick can work for rather small heat fluxes. The degradation was ascribed to the additional thermal resistance of the mesh layers. The association of the meshes and grooves induced nucleation boiling so that the heat pipe performance turned from capillary-limit to boiling-limit. Hsieh et al. [5] tested flat-plate heat pipes with mesh-covered rectangular grooves with a depth of 0.55 mm and a width of 0.5 mm. One, two, or three layers of 70 lm-thick mesh was plated on the grooves with a supportive structure. The working fluid was acetone. The results for the
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Nomenclature Ac Ae Ah Ap d k Leff N Dpc q Q Qcond Qmax Qps Qt Rc Re Re,min rp t T T
cooled area [cm2] heated area [cm2] cross-sectional area of the heating block [cm2] cross-sectional area of the base plate [cm2] mesh wire diameter [mm] thermal conductivity [W/m °C] effective length of the heat pipe [mm] mesh number [1/mm] capillary pressure [Pa] net heat flux [W/cm2] net heat load (= Qt Qcond) [W] lateral conduction through copper plate [W] maximum heat load before dryout [W] heat input from the power supply [W] total heat load [W] condenser resistance [°C-cm2/W] evaporator resistance [°C-cm2/W] minimum evaporator resistance [°C-cm2/W] pore radius of the wick wick thickness [mm] temperature [°C] thermocouple
horizontal orientation showed that the composite wick with a single mesh layer could improve thermal performance marginally at a small heating area but not at a larger area. With two or three mesh layers, the maximum heat transfer rate decreased and the thermal resistance increased. Under tilt angles, although the composite wicks could maintain heat pipe operation, the thermal performance was seriously degraded. In developing a polymer-based flat-plate heat pipe, Oshman et al. [6] fabricated a copper composite mesh-micropillar wick by electroplating a 200-mesh woven screen on top of micropillars. The 200-lm wide inter-pillar cross-shaped channels were used as the liquid passage with low flow resistance. This polymer-based heat pipe was able to operate at a heat flux of 11.94 W/cm2. Oshman et al. [7] subsequently reduced the width of the inter-pillar channels to 31 lm to form sharp corners and promote capillarity. This modified flat-plate heat pipe with a small footprint (40 mm 40 mm) was demonstrated to operate under adverse acceleration up to 10 g. To characterize the enhanced capillary evaporation with visualization, Dai et al. [8] tested a composite mesh-groove wick vertically placed in a chamber under atmospheric pressure. The composite wick was made by sintering one layer of 145-mesh copper woven screen on top of parallel grooves with a 250 lm 250 lm cross-section. The center of the heating area was 15 mm above the water level. The critical heat flux (CHF) for the composite wick was substantially increased over that for a groove or a mesh wick. The primary reasons for the increased CHF were ascribed to the increased capillarity from the mesh pores and reduced flow resistance of the grooves. Dai et al. [9] further investigated the effects of superhydrophilic treatment on the composite wick using the same test apparatus. The heat transfer coefficient was raised due to enhanced thin film evaporation on the mesh. However, the maximum heat flux was not significantly improved as a result of compromise between the enhanced capillarity and increased viscous drag by the coating materials. At low heat fluxes, oscillating flow owing to the growth and collapse of the boiling bubbles was observed, and the heat transfer coefficient was increased by the induced advection. At larger heat fluxes, the channels became partially dry with vapor cores and the capillary force was mainly provided by the sharp mesh-wire corners. The copper meshes provide larger
x y
direction along the heat pipe direction along the heating block
Greek
a d
e h h0
r x
tilt angle [°] thickness of the immersed part of the wick [m] wick porosity apparent contact angle at the evaporating contact lines [°] static contact angle [°] surface tension [N/m] measurement uncertainty
Subscripts a adiabatic section avg average c condenser cu copper e evaporator eff effective v vapor w wick
surface areas and more nucleation sites for heat transfer. The growth and collapse of bubbles intensified the capillary liquid motion on the meshes and enhanced the thin film evaporation thereon. The works of Oshman et al. [6,7] and Dai et al. [8,9] have verified the advantages of the composite mesh-groove wicks. However, the advantages have not been realized in heat pipes by Lefèvre et al. [4] and Hsieh et al. [5]. The failure was likely because the meshes were plated by supportive structures over the grooves without sintering or electroplating. There could be high contact resistances and possible local detachments to deteriorate the capillarity. It is noted that the visualization studies of Dai et al. [6,7] conducted for water were under the atmospheric condition. The nucleate boiling was observed to be significantly influential to the capillary evaporation process. However, under the low pressures in operating heat pipes, whether boiling occurs and other evaporation characteristics for composite mesh-groove wicks needs be observed by visualization experiments. Heat pipe experiments with visualization have been conducted for sintered-mesh wicks [11–15], sintered-powder wicks [16], groove wicks [17–21], and composite sintered-power/groove wicks [2]. For copper/water heat pipes, nucleate boiling was not observed for thin sintered-mesh wicks [11–15], sintered-powder wicks [16], groove wicks [18], and composite sintered-power/groove wicks [2]. When methanol or acetone was used as the working fluid, weak nucleate boiling occurred near the Qmax which appeared to be still limited by wick capillarity [13]. Hong et al. [22] verified that bubble nucleation is harder to occur in a thin (<1 mm) sintered-powder wick at low pressures. Their theoretical analysis showed that, mainly because of the low vapor density, the wall superheats required for onset of nucleate boiling at low pressures are several times larger than at the atmospheric condition. Therefore, visualization experiments are needed to explore the evaporation characteristics of operating heat pipes with a composite mesh-groove wick which is potentially advantageous in both high permeability and high capillarity. The following questions are worthy to be examined: (1) How is the quantitative thermal performance, namely the maximum heat load and the thermal resistance, of this composite wick compared with
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a mesh-wick and a grooved heat pipe? (2) Could its advantages sufficient to maintain heat pipe operation at adverse tilt angles? (3) Would nucleate boiling occur and influence heat pipe performance at the low-pressure conditions in heat pipes? The present work aims to answer these questions with visualization experiments for a copper/water with a sintered composite mesh-groove wick. For comparison, a heat pipe with the composite wick along with a grooved heat pipe and a double-layer meshed heat pipe will be carefully tested the under horizontal orientation. Then, the mesh-wicked heat pipe and the composite-wicked heat pipe will be tested at various tilt angles. The maximum heat transfer rates, the evaporator resistances, and the visualized evaporation characteristics for these three different wicks will be compared. 2. Experimental methods Fig. 1 shows the overall test setup. The evaporation characteristics of a flat-plate heat pipe were visualized through its glass window using a CCD camera equipped with a microscopic lens (Optem, Zoom 125, resolution up to 600 lp/mm) to capture magnified images, with illumination from a couple of LED lights. Fig. 2 shows the detailed construction of the heat pipe and the thermocouple locations. A wicked base plate of the heat pipe is shown in Fig. 3. The wicked region is 126 mm 50 mm, and the effective length (Leff = 0.5 Le + La + 0.5 Lc, where Le, La, and Lc are the length of the evaporator, adiabatic section, and the condenser, respectively) of the heat pipe is 101 mm. The three different wicks tested in the present work are: a parallel-groove wick, a sintered mesh-groove wick, and a sintered double-layer 200 mesh wick. The crosssectional views of the parallel-groove wick and the mesh-groove
841
Fig. 3. The wicked base plate.
wick are shown in Fig. 4. The grooves were chemically etched, by a commercial etching company, on a 4 mm-thick C1020 oxygenfree copper base plate. The grooves have a roughly semi-circular cross-sectional shape with a width of 0.18 mm and depth of 0.09 mm (Fig. 4a). The sintered composite mesh-groove wick was made by sintering a layer of 200 mesh screen on top of the grooves under a load of 1 kgf/cm2 in a sintering oven with an 800 °C hydrogen/ nitrogen atmosphere for 2 h. Before sintering, the grooved base plate and the mesh screens were carefully cleaned with acetone and water for several times. After sintering, the total depth of the composite sintered wick is 0.18 mm, with a 0.09 mm depth of each portion (Fig. 4b). The average porosity of the composite sintered wick is determined by averaging the groove and the mesh parts. The groove part is 0.50, measured as the cross-sectional area fraction of the grooves. The mesh part is determined using the relation adapted from that for a single-layer 200 mesh wick [10]:
Fig. 1. Experimental setup.
Filling tube
Fig. 2. The flat-plate heat pipe with thermocouple layout.
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(a)
(b)
(c)
0.15mm
Fig. 4. Cross-sectional views of (a) the groove wick, (b) the sintered composite mesh-groove wick, and (c) the 2 200 mesh wick.
e ¼ 1 2:1pNd2 =t; where N is the mesh number, d is the wire diameter, and t is the wick thickness after sintering. With N = 7.874/mm, d = 0.055 mm, and t = 0.09 mm, the porosity of the mesh part is 0.56. Therefore, the porosity of the composite mesh-groove wick is 0.53. The double-layer mesh wick (named as 2 200 mesh wick) was made by sintering two layers of 200 mesh screens on a plain base plate under the same sintering condition as for the composite wick. Its wick porosity e is determined using the following relation for a double-layer mesh wick:
e ¼ 1 1:05pNd2 =t: With t = 0.15 mm after sintering, as shown in Fig. 4c, e = 0.48. The static contact angles for various surface conditions were measured right before the assembly of the heat pipe using a contact angle meter (Mactech Co.). The measurements were made for sessile drops of 5 ll at three locations on the flat edge region of the copper base plate. The static contact angles were h0 = 13 ± 1°, in agreement with our previous measurements [2,12–15,18]. The present contact angle values are smaller than most of the literature values [23–25], which scatter between 7° and 94° and are sensitive to surface roughness, cleansing process and oxidation condition. As our copper surfaces were carefully cleaned and reduced in a hightemperature hydrogen/nitrogen atmosphere, low values have been obtained consistently. To keep its h0 within the range of 13 ± 1°, a
reduction process was necessary for each test run. The reduction process was similar to the sintering process except that no pressure was exerted on the base plate. To avoid degradation of the contact angle, evacuation of the assembled heat pipe started no more than 5 min after the base plate was taken out of the sintering oven. Also, after the tests the base plate could maintain good wettability. The inner space of the heat pipe was 126 mm 50 mm 5 mm. With the interfaces between different pieces sealed with orings, the whole structure, including the wicked copper base plate, the upper glass window, and stainless-steel frames, was tightened with bolts. A uniform heat load Qt, through a copper block, was supplied by a cartridge heater embedded in another copper block below the upper one. The cartridge heater was powered by a DC power supply. The heating area was 10 mm 50 mm; the watercooled area was 40 mm 50 mm, provided by a cold plate with cooling water of 20 ± 0.5 °C. The interfaces between the heating block/cold plate and the base plate were laid with a layer of thermal grease (Dow Corning TC5021, thermal conductivity of 3.3 W/ mK). The heating block and the cold plate were insulated within a Bakelite box, on which the flat-plate heat pipe was positioned. The heat pipe and the Bakelite box were altogether wrapped with ceramic wool except for short periods of video recording. Since most of the tests were conducted with the adiabatic-section temperatures near the room temperature, condensation on the glass window was not serious in our experiments as the heat loss through the window was limited. When the window was indeed blocked by a water layer, a hair blower was used to evaporate the water on the window. Afterwards, thermal equilibrium was re-attained before further measurements. To reduce the longitudinal conduction through the 4 mm-thick copper base plate, trenches were made across the plate and around the evaporator and the condenser to leave a local plate thickness of 0.8 mm (Figs. 2 and 3). The fluid charge was somewhat more than the saturate amount to fill the wick voids. The excess charge was needed to compensate for the liquid trapped at the corners, in the gaps, and on the walls of the heat pipe. Nineteen K-type SS-sheathed probe thermocouples (Omega, Inc.) were used to measure the temperatures at selected positions. Among them, these measuring the internal vapor temperatures were inserted via holes drilled through the middle stainless-steel frame. To prevent air leakage, vacuum fittings were installed to accommodate these /1.6-mm thermocouples. All the other probe thermocouples were /1.0 mm. Connected to a data logger (Fluke, Hydra Series II), they yielded temperature readings with a resolution of 0.1 °C. The readings of these thermocouples have been checked to agree well with each other under the room temperature and the boiling water. As shown in Fig. 2, thermocouples T1, T2 and T3, each 5 mm apart, were embedded in the upper heater block to measure Qt. Calculations used T1 and T3 (T2 as a reference for linearity checking) and the relation (Dy = 10 mm)
Q t ¼ kcu Ah ðT 1 T 3 Þ=Dy;
ð1Þ
where kcu is the thermal conductivity of copper, determined from the standard physical property data, and Ah is the cross-sectional area of the heating block. Thermocouples T4–T6 measured temperature T4–T6 of the copper plate in the heated area, with T5 at the center. T7 was positioned at 1 mm above the center of the evaporator to reflect the vapor temperature T7 leaving the evaporator. They were used to determine the evaporator resistance. The lateral conduction through the copper plate, Qcond, was estimated using a pair of thermocouples embedded in the copper base plate (T9 and T10) with
Q cond ¼ kcu Ap ðT 9 T 10 Þ=Dx;
ð2Þ
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where Ap is the cross-sectional area of the base plate and Dx is the distance between thermocouples T9 and T10. In general, these calculated ratios of Qcond/Qt were about 3–5% prior to the onset of partial dryout in the composite-wick evaporator. The net heat load Q was determined by
Q ¼ Q t Q cond :
ð3Þ
Thermocouples T11, T12, and T13, 10 mm apart, measured the base plate temperatures T11, T12, T13 below the condensation zone. All the thermocouples for base-plate temperature measurement (T4–T6 and T9–T13) were positioned along the central axis of the heat pipe. At 1 mm above the condenser wick were T14, T15, and T16, which measured the vapor temperatures needed to calculate the condenser resistance. Furthermore, 4 mm above the condenser wick were placed T17, T18, and T19 to monitor the extent of noncondensable gas (NCG). The NCG might have entered along with the charge water or leak in through the o-ring sealing during the test. To reliably measure the heat pipe performance, NCG was carefully minimized in the charging process and monitored throughout the tests. De-ionized water was used as the working fluid. Before each test, the working fluid was degassed by boiling in a long-neck flask, whose outlet is connected with a Graham condenser, for a sufficient duration. Right after the heat pipe was evacuated down to a pressure of 8 103 Torr, a selected volume of degassed liquid was filled into the heat pipe via a filling tube welded on the stainless-steel frame (Fig. 2). Afterwards, the filling tube was shut by a valve. The heat pipe was then set up for testing. The heating from the DC power supply began from a low power and was incremented stepwise. Except for a special test run, the duration of each test was within 4–6 h. Data were taken under a thermally stable condition for each heat load, with all temperature variations less than 0.2 °C within the last 10 min. Since the dryout behaviors are quite different for the three types of wick, the tests proceeded until the base temperatures at the evaporator increase drastically, reflecting partial or even full dryout in the evaporator. The Qmax was then determined as the maximum heat load before the drastic increase in Re. Since the small amount of water trapped at the corners, gaps and walls varied in different tests, the results of Qmax, Re, and especially the condenser resistance Rc for the same test conditions may vary. The evaporator resistance Re and the condenser resistance Rc were determined respectively as
Re ¼ ðT e;av g T 7 Þ=qe Rcu ½ C-cm2 =W
ð4Þ
Rc ¼ ðT v ;av g T c;av g Þ=qc Rcu ½ C-cm2 =W
ð5Þ
where Te,avg = (T4+2T5+T6)/4, Tv,avg = (T14+T15+T16)/3, Tc,avg = (T11 + T12 + T13)/3, based on the portion of coverage of each thermocouple in the measured region; Rcu (= 0.0875 °C-cm2/W) accounts for the thermal resistance associated with the distance of 3.5 mm between the wick bottom and the thermocouples embedded in the grooves made on the lower surface of the copper plate. Since the heat pipe was well insulated during the tests, the heat loss to the environment was neglected. Hence the heat Q via evaporation was considered equal to the heat via condensation. In Eqs. (4) and (5), the net heat fluxes qe = Q/Ae, and qc = Q/Ac, with the heated area Ae = 5 cm2 and the cooled area Ac = 20 cm2. The experimental uncertainties were calculated using the propagation-of-error method. As the power supply input was calculated by Qps = I V, the relative uncertainty of Qps can be determined as
x Q ps Q ps
" #1=2 xðIÞ 2 xðVÞ 2 ¼ þ : I V
ð6Þ
With x(I) = ±0.005 Amp and x(V) = ±0.05 Volt, the uncertainty of Qps was less than ±1%. The relative uncertainty of Qt was determined according to Eq. (1) as
xðQ t Þ Qt
" ¼
xðkcu Þ kcu
2
þ
xðAh Þ Ah
2
þ
2
xðDTÞ DT
2 #1=2
þ
xðDyÞ
:
Dy
ð7Þ The relative uncertainty for kcu was ±0.5% and those for Ah and Dy were negligible. In temperature measurements, the thermocouple readings were rather stable. The average of these readings lay within an uncertainty of ±0.05 °C for a single thermocouple. Thus, the uncertainty in DT associated with two thermocouples was ±0.07 °C. For the composite wick, the operation range of Q was 13.4–66 W. For 30 W < Q < 66 W, the values of (T1 T3) used in Eq. (1) to calculate Qt ranged between 1.6 and 4.0 °C. Hence, the relative uncertainty in Qt was between ±2 and 5%, with a larger uncertainty at a lower Qt. For 17 W < Q < 30 W, (T1 T3) ranged between 0.9 and 1.6 °C. Hence, the relative uncertainty in Qt was between ±5 and 8%. The maximum uncertainty in Qt was ±10% at the lowest Q of 13.4 W. The relative uncertainty in Q (= Qt Qcond) was near that of Qt since Qcond was only a small fraction of Qt. For the 2 200 mesh wick, the Q range was 12–26 W with (T1 T3) between 0.7 and 1.3 °C. Hence, the uncertainty in Q was between ±5 and 10%. For the groove wick, Q ranged between 9.4 and 14.1 W with (T1 T3) between 0.6 and 1.1 °C, and the uncertainty in Q was between ±6 and 12%. The uncertainties in Re and Rc were calculated using the propagation of error method based on Eqs. (4) and (5), respectively. For the composite wick, the relative uncertainties in Re were ±17% or 26% for Q = 13.4 W (two cases); ±9– 14% for 17 W < Q < 30 W; ±2–9% for 30 W < Q < 66 W, and ±4– 14% in Rc with a larger uncertainty at a lower Q. For the 2 200 mesh wick, the uncertainties were ±7–14% in Re and ±7–14% in Rc. For the groove wick, the uncertainties were ±7–15% (but ± 37% at 9.4 W) in Re and ±9–14% in Rc. The measurement uncertainties for heat pipes with different wicks are summarized in Table 1. Another way to roughly estimate the uncertainty in Qt is by comparing the power supply input and the heat-conduction calculation by Eq. (1). In general, their differences are within 15%. 3. Results and discussion 3.1. The composite mesh-groove wick Fig. 5 illustrates the longitudinal temperature distributions in the vapor space and the copper base plate for the horizontal heat pipe with the composite mesh-groove wick at four representative heat loads. The liquid charge v in this test run is 0.9 ml (the saturate charge to fill the wick is 0.6 ml). In this figure, the vapor temperatures above the condenser are represented by T17–T19
Table 1 Summary of measurement uncertainties for different wicks. Composite mesh-groove wick
2 200 mesh wick
Groove wick
Qps
<±1%
Qt or Q
±10%, Q = 13.4 W ±5–8%, 17 W < Q < 30 W ±2–5%, 30 W < Q < 66 W
±5–10%
±6–12%
Re
±17% or ±26%, Q = 13.4 W ±9–14%, 17 W < Q < 30 W ±2–9%, 30 W < Q < 66 W
±7–14%
±7–15% (but ±37% at 9.4 W)
Rc
±4–14%
±7–14%
±9–14%
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φ1.6 mm TC
Fig. 5. Longitudinal temperature distributions in the vapor and the base plate at different heat loads for the horizontal heat pipe with a sintered composite meshgroove wick, v = 0.9 ml.
partial dryout
heated zone
Fig. 6. Image of the evaporator with partial dryout, v = 0.9 ml, a = 90°, Q = 39.4 W.
2
2 The reasons for the differences between the upper and lower thermocouple readings have been discussed in [15]. 3 As Q is increased to 56.5 W or 63.4 W, the vapor temperatures (T17–T19) become lower than that at the adiabatic section (T8), reflecting mild NCG effect. It is noted that the total test time of this specific test run (Case 1 with v = 0.9 ml shown in Fig. 7), with data taken at a large number of heat loads, was about 8 h. Some amount of NCG might have leaked in during the long test time. All other tests were within 4–6 h, and the NCG were controlled within a limited extent.
Re, case1, v = 0.9 ml Re, case2, v = 0.95 ml
composite mesh-groove wick
Rc, case1 , v = 0.9 ml
1.5
R (K∙cm²/W)
measured by the upper thermocouples T17–T19. Among the six thermocouples (T14–T19) above the condenser (cf. Fig. 2), the lower three (T14–T16) 1 mm above the condenser wick were used for Rc calculation (Eq. (5)), while the upper three (T17–T19) 4 mm above the wick were used to monitor the NCG effect. The present results are in similar trends as those obtained in our previous measurements [14,15].2 At Q = 19.5 W and 42.2 W, the vapor temperatures are rather uniform even near the end of the condenser region, showing negligible NCG accumulation near the condenser end during the early period of the test.3 This feature also evidences the insignificant longitudinal pressure drop and heat loss (either through the upper window or to the base plate) associated with the vapor flow across the upper space of the chamber. At Q = 56.5 W, partial dryout has occurred in the heated zone so that the plate temperatures rise drastically. In addition, the left-end thermocouple reading is the highest, with T4 > T5 > T6 because partial dryout initiates from the end of the evaporator region. The image in Fig. 6 reflects the uneven distribution of the partial-dryout region over the heated zone for the composite mesh-groove wick. When partial dryout occurs in certain grooves, the hot dryout region would prompt the onset of partial dryout in neighboring grooves. Besides, the dryout lengths in individual grooves are different. With the heat load increased, the dryout regions expand until the whole heated zone dries out, yielding high plate temperatures. The condition at Q = 63.4 W presented in Fig. 5 corresponds to the situation when the dryout region almost covers the whole heated zone. At Q = 56.5 W (partial dryout) and 63.4 W (near full dryout), the vapor temperatures measured above the heated zone (T7) are higher than those at the adiabatic section (T8). This is because thermocouple T7 is influenced by the upstream vapor above the dryout region. After sufficient mixing, the vapor temperature settles to the equilibrium temperature of T8. Another point to note is that the slight drop of the plate temperature at the condenser end (T13) is because the cooling water enters the cold plate from the end side. Fig. 7 presents the evaporator and condenser resistances for two horizontal test runs for the composite-wick heat pipe. Fig. 8 shows the magnified images of the evaporator at different heat loads,
Rc, case2, v = 0.95 ml
1
0.5
0 0
10
20
30
Q(W)
40
50
60
70
Fig. 7. Evaporator and condenser resistances versus Q for two test runs for the composite wick.
with the images taken from Case 2 (v = 0.95 ml) in Fig. 7. For both cases, the evaporator resistances (Re) remain around a fixed value up to Q 43 W. In this stage, the evaporator remains fully wetted, as shown in Fig. 8a, thanks to the capillary force provided by the small mesh pores. When Q > 43 W, partial dryout initiates from the upstream end of the heated zone. The dryout region expands with increasing Q, leading to gradually increased Re, as shown in Case 2. Fig. 8b shows two neighboring grooves with one dried and another wetted. During this stage, the values of Re for Case 1 remains near a constant lower than for Case 2. But for a third run not shown here, the variation trend of Re is similar to that of Case 2 with gradually rising Re during the partial-dryout stage. The reason why the Res of Case 1 are distinctly lower than those of the other two cases is unclear. When Q is further increased beyond 60 W or 63 W, another jump in Re appears. Fig. 8c shows the image taken around the downstream edge of the heated zone at Q = 66 W, when the dryout region almost covers the whole heated zone. At this heat load, Re is stably maintained at a high value. Note that at this test condition, Qcond/Qt increases to 12% as a result of the high evaporator temperature. In comparison, Qcond/Qt remains less than 5% in the fully-wetted stage and between 5 and 9% in the partial-dryout stage. As far as the condenser resistances Rc are concerned, the present values are in similar trend as our previous experimental results [14,15] that Rcs are several times larger than Res. Note that the Rc data are sensitive to the amount of liquid trapped in the internal gaps or on the walls of the heat pipe. Since the trapped amount varied in different tests and different times in the same test, the Rc data can only be regarded as qualitative reference.
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Fig. 8. Magnified images of the evaporator, (a) fully wetted at Q = 29.8 W, (b) partial dryout at Q = 47.5 W, (c) near full dryout at the downstream edge of the heated zone, Q = 66.0 W.
Fig. 9 compares the evaporator resistances for a composite-wick heat pipe under various tilt angles. All the cases, horizontal or tilted, exhibit nearly unvaried Res over the first stage up to 39– 49 W. During the first stage, the heated zone is largely fully wetted. This is because when the grooves are covered with a mesh layer, they function as capillary tubes with the mesh pores providing intense capillary force. Thus, dryout within the grooves is postponed and good anti-gravity ability can be gained. Even at a high tilt angle of 90°, the low-Re stage can be maintained up to 39.4 W. Beyond the first stage, different Re–Q trends are observed for the two a = 30° cases. In Case 1, Re rises gradually as Q > 40 W, while in Case 2, Re remains low up to Q = 49 W, followed by a drastic jump. The gradual Re rise in the former case is similar to Case 2 of a = 0°, during which partial dryout expands gradually over the
1
Re ( C∙cm /W)
φ1.6 mm TC
0° case 1 , v = 0.9 ml 0° case 2 , v = 0.95 ml 30° case 1 , v = 0.9 ml 30 ° case 2, v = 0.9 ml 45 ° , v = 0.95 ml 90 °, v = 0.9 ml
0.8
0.6
heated zone. Drastic jump at the end of the first stage occurs for a 45°. The partial-dryout stage, which sustains substantially at a = 0°, stays within a narrow range of heat load. The narrow transitional partial-dryout stage is illustrated using Figs. 6 and 10 for a = 90°. In Fig. 6, at Q = 39.4 W when partial dryout has first appeared, about 1/3 of the heated zone has dried out. But when Q is increased to 43.4 W, the dryout region expands to cover almost the whole heated zone, as shown in Fig. 10. Note that the elapse time for the expansion to stabilize was as long as 40 min. For high tilt angles, once partial dryout appears, it could develop to full dryout of the heated zone after a mild increase of heat load, leading to a jump in Re. In contrast, the dryout expansion process for a = 0° was gradual and the elapse times for stabilization were about half as long as for a = 90°.
0.4
dryout
heated zone
wetted
0.2
0
0
10
20
30
40
50
60
70
Q (W) Fig. 9. Evaporator resistances versus Q for the composite wick under various tilt angles.
Fig. 10. Image of the evaporator at a = 90° with the heated zone nearly fully dried at Q = 43.4 W.
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3.2. The groove wick Fig. 11 illustrates the variations of Re and Rc versus Q for a horizontal grooved heat pipe with v = 0.6 ml (the saturate charge to fill the groove volume is 0.3 ml). At Q = 9.4 W, the groove wick is fully wetted, as shown in Fig. 12. When Q is increased to 11.6 W, partial dryout can be observed. At Q = 13.2 W, the dryout region has extended over the middle line (marked by the thermocouple) of the heated zone, as shown in Fig. 13. When Q = 14.1 W, full dryout has prevailed the heated zone. The zigzag shape of the steepsloped liquid fronts has been observed by Wong and Chen [18].
1.6 groove wick, v = 0.6 ml
Re
1.4
The back-and-forth oscillation of the liquid fronts observed in [18] was not obvious in the present tests. For the case shown in Fig. 11, the liquid fronts were stagnant at a fixed Q. In some other tests, very mild oscillations were observed. The liquid frontline receded with increasing Q, leading to higher temperatures in the heated zone and hence higher Re. Detailed description of the receding process of the liquid fronts in connection with the continuous temperature increase in the heated zone has been provided in [18]. No boiling was observed, the same as in Wong and Chen for water [18]. For grooved heat pipes, Re rises continuously with increasing Q as a result of continuous liquid-front recession. This has been indicated by the visualization experiments of Wong and Chen [18] and by Lin and Wong [26] for commercial grooved heat pipes. The groove wick cannot operate under tilted condition with a 30°.
Rc
R ( C∙cm²/W)
1.2
3.3. The 2 200 mesh wick
1 0.8 full dryout 0.6 partial dryout 0.4 0.2 0
fully wetted 0
5
10
15
Q (W)
20
Fig. 11. Evaporator resistances versus Q for the horizontal heat pipe with the groove wick.
heated region
Fig. 14 shows the Re and Rc results versus Q for a horizontal heat pipe with the 2 200 mesh wick. The Re data of the four runs are quite consistent. The Qmaxs range around 21–25 W, followed by the onset of dryout and marked increase in Re. The minimum evaporator resistance Re,min occurs around Qmax, as a result of thinnest liquid layer thickness in the evaporator. Detailed description of the dryout process for multi-layer mesh wicks in connection with visualization images and temperature measurements at the evaporator is available in Liou et al. [11] and Wong et al. [13]. Similar as in Liou et al. [11], no boiling was observed here for water. The values of Rc at v = 1.3 ml (the saturate charge is 0.5 ml) are higher than those at v = 1.1 ml due to the thicker liquid layer at the condenser. The condensation mechanisms for a mesh wick with different surface wettabilities, rendered by different working fluids or surface treatments, have been carefully discussed by Wong et al. [14,15]. Tests under a = 30° and 60° were attempted, but the heated zone dried out at Q = 14 W, yielding high Re values as large as 0.66 °C-cm2/W and 1.0 °C-cm2/W, respectively. Under a tilt angle, the wick capillary was insufficient to sustain wetted evaporator against liquid flow resistance and additional gravity at this heat load. 3.4. Comparison between different wicks
thermocouple
Fig. 12. Image of the fully wetted groove evaporator, Q = 9.4 W.
centerline of heated region
liquid front
According to the above heat pipe test results, the sintered composite mesh-groove wick is remarkably superior to the others, due to the combinative advantages of high capillarity from the mesh screen and high permeability of the parallel grooves. For horizontal operation of the composite wick, Qmax may be defined as the
1.8 Re, v = 1.1 ml, case 1
1.6
1.4
R ( C∙cm²/W)
2 200 mesh wick
Rc, v = 1.1 ml, case 1 Re, v = 1.1 ml, case 2
1.2
Rc, v = 1.1 ml, case 2
1
Re, v = 1.3 ml, case 1
Rc, v = 1.3 ml, case 1
0.8
Re, v = 1.3 ml, case 2
0.6
Rc, v = 1.3 ml, case 2
0.4 0.2
1.6 mm wetted
dryout
thermocouple
Fig. 13. Image of the groove evaporator with partial dryout, Q = 13.2 W.
0
0
5
10
15
Q (W)
20
25
30
Fig. 14. Evaporator and condenser resistances versus Q for the horizontal heat pipe with the 2 200 sintered mesh wick.
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maximum Q prior to the full dryout, since Res remain reasonably low. Then, the Qmax of about 60 W is much higher than 21–25 W for the 2 200 mesh wick and about 10 W for the groove wick. Even under a tilt angle of 30–90°, the Qmax may reach 39–49 W. In contrast, the 2 200 mesh wick cannot operate at Q = 14 W under a = 30°. The negative evaluation on the composite mesh-groove wick of Lefèvre et al. [4] was likely because the meshes were plated over the grooves only by supportive structures. There would be high contact resistances and possible local detachments to deteriorate the capillarity. However, in the present study and [6–9], in which the mesh screen was sintered or electroplated onto the grooves, positive results have been obtained. Therefore, secure bonding between the mesh screens and the grooves, either by sintering [8,9] or electroplating [6,7], is essential for the composite meshgroove wicks to manifest their superiority as the heat pipe wick. 4. Conclusions Visualization experiments have been performed on flat-plate heat pipes with a sintered composite copper mesh-groove wick at different tilt angles. For comparison, a groove wick with the same groove size and a sintered 2 200 mesh wick also have been investigated. The effective length of the heat pipes is 101 mm. Deionized water is used as the working fluid. The following conclusions are reached: 1. The composite wick exhibits two evaporation stages under the horizontal orientation before full dryout in the heated zone. In the first stage with Q up to about 40 W, the wick is filled with water; in the second stage, partial dryout in the grooves occurs and expands with increasing heat load. Up to about 60 W, the heated zone becomes fully dried. In contrast, the Qmax of the 2 200 mesh wick is 21–25 W, and full dryout prevails at 14 W for the groove wick. 2. When the tilt angle is between 30° and 90°, the Qmax for the composite wick may reach 39–49 W, followed by drastic increase in the evaporator resistance. The 2 200 mesh wick suffers serious dryout at Q = 14 W under a = 30°; the groove wick cannot operate under a 30°. 3. The sintered composite mesh-groove wick is superior to the multi-layer mesh wick and the groove wick because of its combinative advantages of high capillarity from the mesh screen and high permeability of the parallel grooves. 4. No nucleate boiling is observed in all the present tests for the different wicks with water as the working fluid. The heat pipe performance is capillary-limited. 5. Secure bonding between the mesh screens and the grooves, either by sintering or electroplating, is essential for the composite mesh-groove wicks to manifest their superiority as the heat pipe wick. The present performance tests have verified the superiority of the sintered composite mesh-groove wick. Our future work will measure the permeabilities and the effective pore sizes of these different wicks. Working fluid other than water will also be tested. Until then, the superiority of the sintered composite mesh-groove wick as a result of its combinative advantages of both high capillarity and high permeability can be quantitatively proved. Disclosure statement All authors of this paper have no actual or potential conflict of interest including any financial, personal or other relationships with other people or organizations.
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