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Spray cooling with ammonium hydroxide
International Journal of Heat and Mass Transfer 107 (2017) 45–52
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Spray cooling with ammonium hydroxide Huseyin Bostanci ⇑, Bin He, Louis C. Chow Department of Mechanical and Aerospace Engineering, University of Central Florida, Orlando, FL 32816, United States
a r t i c l e
i n f o
Article history: Received 26 April 2016 Received in revised form 8 November 2016 Accepted 10 November 2016
Keywords: Thermal management Spray cooling High heat flux Binary mixture Ammonium hydroxide Diffusion resistance
a b s t r a c t An experimental study was conducted to investigate the performance characteristics of spray cooling with ammonium hydroxide (NH4OH) binary mixture for high heat flux removal. Two mixtures, having ammonia mass fractions of 0.3 and 0.5, were selected to represent practical operation conditions near atmospheric pressure and room temperature. Experimental setup involved a closed loop system with a vapor atomized spray nozzle and a 1-cm2 heater sample that simulated a high heat flux source. Tests were performed with gradually increasing heat fluxes of up to 800 W/cm2 and maintaining surface temperatures below 75 °C at varying liquid and vapor flow rates. Results indicated that the heat transfer coefficient (HTC) values from NH4OH mixtures can be lower than those from pure water and pure ammonia. The data suggested that boiling depression, due to mass diffusion resistance at liquid vapor interface, could greatly affect the overall spray cooling performance, especially when the binary mixtures comprise components with widely different boiling points. The study therefore provides performance characteristics, as well as some fundamental insights, for a potential spray cooling scheme suitable for low temperature, low pressure operations in various applications including thermal management of aerospace electronics and electro-optics. Ó 2016 Elsevier Ltd. All rights reserved.
1. Introduction Two-phase spray cooling has been an emerging thermal management technique characterized by its major capabilities of high heat transfer rate, near-uniform surface temperature, and efficient coolant usage that leads to compact and lightweight systems. Due to these capabilities, spray cooling is a promising approach for high heat flux applications in computing, power electronics, and electro-optics. Selection of working fluid is very important as it directly affects both performance characteristics and operation conditions of a cooling system. The working fluid should have a high latent heat of vaporization to remove high heat flux levels, a boiling point close to the desired operating temperature of the target device being cooled, and a freezing point low enough to provide protection against system freeze-up at anticipated cold ambient. A proper temperature-pressure relationship is also critical to avoid excessively low pressure (high vacuum) or excessively high pressure in the system. Moreover, the working fluid should be chemically stable, noncorrosive, and safe.
⇑ Corresponding author at: University of North Texas, Department of Engineering Technology, UNT Discovery Park, 3940 North Elm St. F115, Denton, TX 76207, United States. E-mail address: [email protected] (H. Bostanci). http://dx.doi.org/10.1016/j.ijheatmasstransfer.2016.11.035 0017-9310/Ó 2016 Elsevier Ltd. All rights reserved.
Certain applications, such as thermal management of highpower diode-pumped solid-state lasers for aerospace systems, poses multiple challenges and necessitates careful consideration to determine a proper working fluid. The laser diode arrays generate high heat fluxes at substrate-level (>600 W/cm2), they mandate operation at low temperatures (<25 °C) to attain high electrical-tooptical energy conversion efficiency, and the emitters (and surface temperature of diode arrays) also require nearly-uniform temperature distributions, within a few °C, to achieve high quality optical beam. Due to the ambient conditions for aerospace applications, the cooling system should be able to operate at very low temperatures (70 °C) without freezing issues. Additionally, the system pressure should ideally be close to atmospheric pressure levels during operation. Based on all these challenging conditions, two-phase spray cooling seems to be the appropriate thermal management scheme. Considering thermophysical properties, particularly high latent heat of vaporization, water and ammonia are the two best heat transfer fluids. However, each fluid has undesirable saturation temperature-pressure characteristics for diode laser cooling. In order to use water for low temperature cooling, its boiling point should be reduced significantly from 100 °C at 1 bar to 20 °C at 0.02 bar by maintaining the system under high vacuum. This would then complicate the system design, and cause potential leak issues. Water also has a very high freezing point for aerospace con-
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Nomenclature A CHF h hfg HTC I k l
area, cm2 critical heat flux, W/cm2 heat transfer coefficient, kW/m2 °C latent heat of vaporization, kJ/kg heat transfer coefficient, kW/m2 °C current, A heater wall thermal conductivity, W/m °C TC to spray surface distance in heater wall, m
ditions. Ammonia appears to be an ideal working fluid for spray cooling of diode laser arrays due to its high latent heat, low boiling point and low freezing point at the normal pressure. However, ammonia has a very strong temperature-pressure dependency. The ammonia pressure can easily exceed 10 bar at room temperature. High saturation pressure (4–10 bars) would deform the heat source structure (i.e., diode bar package) and may lead to failure. This is especially true when the substrate accommodating the heat source is made very thin in order to minimize the internal thermal resistance. Therefore, for room temperature spray cooling applications, water dictates very low system pressure leading to serious fluid sealing problems, while ammonia dictates very high system pressure leading to serious structural problems. In this case, utilizing a mixture of compatible fluids is an attractive approach to obtain the desirable saturation temperature-pressure behavior needed for spray cooling of diode lasers. Ammonium hydroxide (NH4OH) is made by dissolving ammonia into water. Varying the mass fraction of the solution can modify the saturation temperature-pressure relationship. For instance, NH4OH with an ammonia mass fraction, x1, of 0.45 provides 0 °C saturation temperature at approximately 1 bar pressure. Its latent heat of vaporization is still 5–10 higher than those of common refrigerants. The low freezing point (around 80 °C for x1 >0.3) also prevents the plumbing from freezing when spray systems are used in aerospace vehicles. This binary fluid takes advantage of the high latent heat of vaporization associated with the parent fluids without the drawbacks of excessively low or high saturation pressures of water and ammonia, respectively. In addition, NH4OH has a reasonably high thermal conductivity comparable to that of water. Both the parent fluids of NH4OH have very favorable spray cooling heat transfer rates as reported in the earlier research [1–6]. If the binary mixture of NH4OH can provide cooling characteristics somewhere between those of its parent fluids – water and ammonia, it would enable spray cooling technology for high heat flux cooling capability at room temperature operation. This would then allow effective cooling of the delicate electronics and optics devices, such as diode laser arrays, without the excessively high or low system pressure. Binary mixtures have been studied extensively for boiling heat transfer due to their wide use in industrial applications, such as petro-chemical, refrigeration, and power generation. Carey [7] provided a comprehensive review on the pool boiling of binary mixtures addressing thermodynamics aspects and major heat transfer characteristics, specifically the effects of mass diffusion resistance on nucleate boiling, and surface tension gradients on CHF. Some of the exemplary efforts in this area focused on developing prediction methods, including Thome and Shakir [8], Fujita et al. [9], and Kandlikar [10] regarding HTC, and McGillis and Carey [11], Fujita and Bai [12], and Yagov [13] regarding critical heat flux (CHF) correlations. However, there are very limited spray cooling research focusing on binary mixtures, as previous efforts mainly
q00 Tsurf TCavg V V_ V_ req x1
q
heat flux, W/cm2 surface temperature, °C average thermocouple reading, °C voltage, V supplied liquid flow rate (ml/cm2.s) required liquid flow rate (ml/cm2.s) liquid mass fraction of more volatile (NH3) component liquid density (kg/m3)
considered single-component working fluids [1–6,14–17], and none of which, to the best of authors’ knowledge, studied NH4OH. Lin et al. [18] investigated spray cooling with water/methanol mixtures at methanol concentrations of 0, 20, 50 and 100% by volume. Their experimental conditions involved various chamber pressures to maintain saturation temperatures of 45 and 65 °C, and relatively low heat fluxes in the range of 15–75 W/cm2. The data showed that HTCs were highest for water, and lowest for methanol. The performance of water/methanol mixtures was found to be between those of pure water and methanol, although in some conditions, the mixture with 50% ethanol concentration offered comparable performance to that of pure water. The authors noted that the nucleate boiling is a major heat transfer mode in spray cooling, and flashing of the more volatile component at the nozzle exit helps with atomizing liquid droplets and enhance the heat transfer performance. Turek et al. [19] evaluated the capabilities of spray cooling for power inverter applications using water/ propylene glycol (WPG) at a 50% concentration by volume. They replaced the single-phase liquid cooled base plate of a commercial power inverter module with a custom-made base plate of comparable size, and integrated a spray nozzle array. DBC board in the module was sprayed with pressure atomizer nozzles featuring 0.19 l/cm2min liquid flow rate and 275 kPa pressure drop. Using WPG at 100 °C, heat fluxes of up to 148 W/cm2 were removed without exceeding Insulated Gate Bipolar Transistor (IGBT) junction temperature of 125 °C. Their spray cooling approach reduced the overall thermal resistance with the help of a very high HTC and achieved up to a 3.4 increase in inverter power level. Later, Turek et al. [20] conducted experiments with pure water at 95 °C using the same setup. IGBT junction temperatures for the water spray cooling was about 10 °C lower compared to WPG spray cooling, although the difference in the boiling points of two cases was 6 °C. Considering the similar subcooling levels, they concluded that the vapor generated by WPG boiling is nearly all water, leaving a higher concentration mixture at the spray surface with a higher boiling. This locally higher boiling point diminishes the boiling performance and leads to higher temperatures. Bostanci et al. [21] evaluated a two-phase spray cooling scheme for the thermal management of automotive power inverter modules using an antifreeze coolant (water/alcohol mixture). The experimental setup included pressure atomized spray nozzles with 0.15 l/cm2min liquid flow rate and 145 kPa pressure drop, and featured two types of enhanced spray surface with microscale structures. The spray cooling tests with azeotropic mixture at 88 °C saturation temperature and atmospheric pressure, showed that up to 400 W/cm2 heat fluxes can be reached with only 14 °C surface superheat, resulting in a HTC of 280,000 W/m2 °C. Compared to single-phase convective cooling, this spray cooling system offered up to 5.5 potential increase in inverter power. Most recently, Karpov et al. [22] performed an experimental study with water/ethanol mixtures, at ethanol concentrations of 0–96% by mass, and used air atomized
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spray nozzle arrays in a pulsed mode resulting in very low flow rates of up to 0.0173 g/cm2.s. They examined the effects of spray duty cycle and frequency with an isothermal heater at 70 °C, and observed increasing HTC values with increasing ethanol concentration (with a max. at 50–60%). The higher HTC was explained with the accelerated evaporation rates driven by the ethanol content. This study aims to experimentally investigate the performance characteristics of spray cooling with NH4OH binary mixture for high heat flux removal at practical operation conditions near atmospheric pressure and room temperature. The initial data will help understand the heat transfer mechanisms for binary mixture spray cooling, and be instrumental in the development of a high heat flux cooling technique for low temperature, low pressure operations in various applications including thermal management of aerospace electronics and electro-optics. 2. Experimental setup and procedure 2.1. Working fluid Ammonium hydroxide (NH4OH) was selected as the working fluid in this study. Aqua ammonia, aqueous ammonia and ammonium hydroxide are synonymous terms referring to a solution of ammonia in water. Achieving a high heat flux cooling technique at low temperature and pressure operation conditions requires to use a working fluid with proper saturation temperature-pressure relations. Fig. 1 includes such saturation temperature-pressure values for water, ammonia and two other NH4OH mixtures. Defining the desired system conditions to be 1.0–1.25 bar pressure, and 0– 30 °C temperature (shown with shaded areas in the plot), NH4OH mixtures at ammonia mass fractions (x1) of 0.3 and 0.5 emerge as proper working fluids. Specifically, NH4OH at x1 = 0.3 has a boiling point of 27 °C under 1 bar, (that would allow implementing thin structures with low thermal resistance due to pressure level), and NH4OH at x1 = 0.5 has a boiling point of 0 °C under about 1.25 bar (that would allow high electrical-to-optical energy conversion efficiencies due to low temperature level). These two NH4OH mixtures would also provide good freezing protection. Fig. 2a shows the freezing point (along with boiling point) of NH4OH as a function of x1 at atmospheric pressure, where NH4OH at
Fig. 2. Boiling and freezing points (a), and latent heat of vaporization (b) of NH4OH as a function of x1 (based on [23,24]).
x1 = 0.3 and x1 = 0.5 have a freezing point of 680 °C. Fig. 2b includes latent heat of vaporization for the NH4OH mixtures, where it sharply drops beyond x1 = 0.1. Table 1 summarizes the critical thermophysical properties of the selected NH4OH binary mixtures, their parent (pure component) fluids, and some other common refrigerants. When compared to HFC-134a and FC-72, NH4OH at x1 = 0.3 and x1 = 0.5 possess much higher latent heat (7–18) and thermal conductivity (6–10) values. Based on the outlined considerations, NH4OH at x1 = 0.3 and x1 = 0.5 were selected for the present study. While the NH4OH at x1 = 0.3 is commercially available, NH4OH at x1 = 0.5 is not commonly offered by chemical vendors. Therefore, NH4OH at x1 = 0.5 was prepared on site by pulling vacuum on an empty storage cylinder, and then carefully filling it with equal masses of distilled water and anhydrous ammonia. NH4OH is corrosive to copper, copper containing alloys (such as brass), zinc, cadmium, silver, aluminum alloys and galvanized surfaces. Thus, the use of NH4OH as the working fluid requires special attention in materials selection during the system design. Furthermore, such cooling systems should be sealed well and located in non-confined areas, as NH4OH leaks would pose health hazards. 2.2. Spray cooling system
Fig. 1. Saturation temperature and pressure relationship for water, ammonia, and NH4OH mixtures (based on [23]).
A closed loop spray cooling system was used to run the experiments. Schematic diagram in Fig. 3 illustrates the three main sections of the system: NH4OH liquid loop, NH4OH vapor loop, and HFC-134a loop. A reservoir, measuring 10 cm in diameter and 19 cm in length, forms the center of the system containing liquid and vapor phases
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Table 1 Thermophysical properties of selected working fluids (based on [23–26]). Properties at 1.013 bar
Water
Ammonia
NH4OH (x1 = 0.3)
NH4OH (x1 = 0.5)
FC-72
HFC-134a
Freezing point (°C) Boiling point (°C) Latent heat of vaporization (kJ/kg) Thermal conductivity (W/m °C) at 25 °C
0 100 2257 0.614
77.7 33.3 1371 0.466
82.7 27.2 1474 0.548
80.1 4.2 1560 0.526
90 56 88 0.057
103.3 26.1 217.2 0.082
and vapor loops were placed inside a fume hood to ensure the NH3 concentration was always at safe levels (<200 ppm) even in the case of leaks, or during the maintenance of the setup. All wetted parts of the system featured NH4OH compatible materials including stainless steel (chamber), PTFE (tubing), polypropylene (fittings), and acrylic (chamber window). The test setup was equipped with thermocouple (TC) probes, pressure transducers, and a computer controlled data acquisition system for accurate data recording. 2.3. Spray nozzles
Fig. 3. Schematic diagram of the spray cooling setup.
of NH4OH, a test heater, and spray nozzles (a main nozzle for cooling the test heater, and four auxiliary nozzles for condensing the generated vapor). NH4OH liquid loop involves a gear pump that takes saturated NH4OH liquid from the bottom of the reservoir and provides a pressure increase. Part of this liquid flow was supplied to the spray nozzle at a controlled rate. The rest of the liquid flow was routed to a heat exchanger to lower the liquid temperature, and the resulting subcooled liquid was then distributed to the four auxiliary nozzles mounted symmetrically onto the reservoir wall near the top to condense the generated vapor and control the spray chamber pressure/temperature. Use of these side sprays as the condenser allowed a greatly expanded heat exchange area involving surface area of the spray droplets, and helped maintain the mixture concentration level in the liquid pool. NH4OH vapor loop involves a diaphragm compressor that takes saturated NH4OH vapor from the upper section of the reservoir and provides the required pressure lift. The vapor flow at the adjusted rate is then supplied to the main nozzle, where it mixes with the liquid stream and facilitates spray atomization. Additionally, the experimental setup has a HFC-134a (a hydrochlorofluorocarbon compound) refrigerant loop that interacts with the NH4OH liquid loop through the heat exchanger. This loop is a typical vapor compression cycle that absorbs heat from NH4OH liquid loop and rejects it to the circulating tap water. Spray chamber temperature and pressure was controlled in two ways: (i) by adjusting the NH4OH liquid flow rate that goes into the auxiliary nozzles, and (ii) by adjusting the temperature of the heat exchanger through low side pressure of the HFC-134a loop. The reservoir also had a port for charging/ discharging the working fluid. The reservoir, and the NH4OH liquid
A compact vapor atomized spray nozzle (Fig. 4a), manufactured by RINI Technologies, Inc., was used as the main nozzle for cooling the test heater. In these nozzles a fine liquid stream is injected into a high velocity vapor stream. The shear force created by the vapor stream atomizes the liquid into fine droplets that are ejected from the nozzle orifice. The nozzle produces a full-cone spray with an angle of 50° that requires a nozzle-to-surface distance of 10 mm to adequately cover the spray surface. The uniformity of droplet density for this nozzle type was found to be within 20% in radial direction from center to edge. The pressure and temperature of the liquid and vapor supplies, as well as the spray chamber were measured to determine the driving pressures across the nozzle, and saturation condition of the coolant. The other four auxiliary nozzles, used for condensing the vapor, were pressure atomized nozzles featuring a simpler design that requires only liquid supply. All the nozzles were made out of stainless steel. 2.4. Heater and spray surface The test heater design is shown in Fig. 4b–c. One side of the heater had a 1 cm 1 cm protrusion (Fig. 4b) where a matching size (1-cm2) thick film resistor (a flip chip type resistor with backpad metallization) was soldered to simulate a heat generating device. In an actual cooling system configuration, a heat-generating device is usually attached to a substrate or an interface layer that is exposed to a spray/cooling medium, and the interfacial thermal contact resistance is inevitable. Thus, it is very important to minimize the resistance for low impact on the device temperature. In this study, the resistor was carefully soldered onto the heater to ensure that the solder layer was thin and free of air pockets. In terms of temperature measurements in the experiments, the contact resistance has no impact as the measurements are away from the interface. The opposite side of the heater featured a similar size (0.9 cm 0.9 cm) recessed area (Fig. 4c) resulting in a 2 mm thick copper substrate under the heat generating resistor, and it was exposed to spray from the main nozzle. The heater was made of copper (C101 grade), a material not compatible with ammonia. In order to provide corrosion protection, spray side of the test heater was coated with a thin (40–60 lm) layer of tin-lead solder resulting in a relatively smooth surface. The heat flux, q00 , was determined from the total power supplied into the thick film resistor per unit area, and it is defined as q00 = V I/A, where V is the DC voltage applied to the resistor, I is the resulting current for 10 Ohm resistance, and A is the resistor base. The heater temperature was monitored with two type-T TCs
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Fig. 4. Vapor atomized spray nozzle (a), and test section (heater) resistor side (b) and spray side (c).
embedded at the mid-section in the heater wall (1 mm below the surface). One of the TCs was located at the center, while the other TC was located 1 mm away from the center as illustrated in Fig. 4b. Temperature difference between the two thermocouple readings did not exceed ±1 °C at 200 W/cm2, and ±2.0 °C at 500 W/cm2 indicating uniform cooling across the surface. The spray surface temperature, Tsurf, was calculated with Fourier’s 1-D steady-state heat conduction equation, and is defined as Tsurf = TCavg (q00 l)/k, where TCavg is the average of two TC readings, l is the distance from TC location to the surface (taken as 1 mm for copper substrate and 50 lm for tin-lead coating layer), and k is the thermal conductivity of the copper and tin-lead solder. 2.5. Test conditions and procedure Spray cooling setup was maintained at saturation conditions of approximately 27 °C temperature and 1 bar pressure for NH4OH tests at x1 = 0.3, and approximately 0 °C and 1.25 bar pressure for NH4OH tests at x1 = 0.5. Cooling curves, in the form of surface temperature (Tsurf) vs. heat flux (q00 ), were generated for each test by increasing the heat flux gradually in steps of 45–65 W/cm2 until the surface temperature reaches approximately 75 °C. This temperature limit was determined considering practical operation of solid state lasers, which rely on low temperatures for achieving high conversion efficiency. At each heat flux level, the resulting heater temperatures were recorded every 2 s over 3 min durations. Major experimental parameters in this study were the liquid flow rate and atomizing vapor pressure (or vapor pressure drop across the nozzle that is proportional to vapor flow rate) to investigate their effects on the spray cooling heat transfer performance. For NH4OH tests at x1 = 0.3, liquid flow rates (per unit area) and atomizing vapor pressures were varied between 1.05–4.20 ml/cm2.s, and 69–179 kPa (10–26 psi), respectively. For NH4OH tests at x1 = 0.5, liquid flow rates and atomizing vapor pressures were selected to be 1.05–2.10 ml/cm2.s, and 69–138 kPa (10–20 psi), respectively. These parameters resulted in a total of 14 experiments as described in Table 2.
Table 2 Test descriptions. Test ID
NH4OH, x1
Liquid flow rate (ml/cm2.s)
Atomizing vapor pressure (kPa)
1 2 3 4 5 6 7 8 9 10
0.3
1.05 2.10 3.15 1.05 2.10 3.15 1.05 2.10 3.15 4.20
69
11 12 13 14
0.5
1.05 2.10 1.05 2.10
138
179
69 138
film resistor to heater body via conduction was estimated to be approximately 4% based on finite element analysis results. Heat loss from the thick film resistor to the ambient environment was negligibly small (<1 W) based on calculations considering natural convection and black body radiation from 100 °C heater surface to 20 °C stagnant air.
3. Results and discussion Tests described in Table 2 were conducted to investigate the spray cooling performance with NH4OH binary mixture at two concentration levels and at varying liquid and vapor flow rates. Fig. 5
2.6. Uncertainty analysis Experimental uncertainties were estimated for all measurements that are critical in performance evaluation. Error involved in the heat flux measurement (considering variations in voltage, current, and area) was ±1.3% at 500 W/cm2. Error in the temperature measurements from the embedded TCs in the heater wall was calibrated to be ±0.2 °C. Spray surface temperature had an uncertainty of approximately ±0.5 °C involving uncertainty in temperature extrapolation from the TC hole to the heater surface at 500 W/cm2. The HTC calculation included ±3.4% uncertainty at 500 W/cm2. Flow rate measurements had ±5% uncertainty due to variable area flow meter characteristics. Heat loss from the thick
Fig. 5. Spray cooling performance with NH4OH at x1 = 0.3 – effect of liquid flow rate and atomizing vapor pressure.
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includes all data from NH4OH tests at x1 = 0.3 representing 27 °C saturation temperature at 1 bar pressure. The change in saturation temperature during these tests (at heat fluxes of >0 W/cm2), due to fluctuations in chamber pressure, was maintained within 3–4 °C range. Heat fluxes were increased gradually until surface temperature reaches up to 75 °C, and the highest heat flux (but, not CHF) was 800 W/cm2 at 3.15–4.20 ml/cm2.s liquid flow rate and 179 kPa atomizing vapor pressure. Each data set was fitted with a second order polynomial curve to better represent and compare the heat transfer performance. When the effect of liquid flow rate (at a given atomizing vapor pressure) is examined (corresponding to same colored data sets with varying shapes), it is obvious that increasing liquid flow rate provides a better heat transfer, as expected, since higher liquid flow rate would lead to higher droplet velocity and higher droplet flux. However, the level of improvement diminishes beyond 2.10 ml/cm2.s suggesting an optimum level. This observation is consistent for atomizing vapor pressures at 69, 138, and 179 kPa. When the effect of atomizing vapor pressure (at a given liquid flow rate) is considered (corresponding to data sets with similar shapes and different colors), increasing atomizing vapor pressure also improves cooling performance, again as expected, since higher atomizing vapor pressure would lead to higher droplet velocity, and smaller droplet sizes. This trend is consistent at 1.05, 2.10, and 3.15 ml/cm2.s liquid flow rates. Fig. 6 presents data from the NH4OH tests at x1 = 0.5 representing 0 °C saturation temperature at 1.25 bar pressure. Similarly, the change in saturation temperature, due to fluctuations in chamber pressure, was limited to 3–4 °C range. The highest heat flux attained at a surface temperature of 75 °C was approximately 500 W/cm2. Increasing the liquid flow rate at a given atomizing vapor pressure at this concentration level resulted in nearly the same cooling performance. Increasing the atomizing vapor pressure at a given liquid flow rate however, helped improve the cooling performance. The effects of liquid flow rate and atomizing vapor pressure on cooling performance are also illustrated in Fig. 7 to offer an easier comparison and specify the heat removal capability at a fixed 65 °C surface temperature for various conditions with both NH4OH mixtures at x1 = 0.3 and x1 = 0.5. Since the liquid and vapor streams are supplied in saturated condition, the heat absorption on the spray surface is attained mainly by the latent heat during liquid vapor phase change. In this case, vaporization rate can be estimated by calculating the required liquid flow rate to absorb a given heat flux over the actual liquid flow rate used, and it can be defined as _ where q is the NH4OH liquid density, hfg V_ req =V_ ¼ ðq00 =ðq hfg ÞÞ=V,
Fig. 6. Spray cooling performance with NH4OH at x1 = 0.5 – effect of liquid flow rate and atomizing vapor pressure.
Fig. 7. Heat removal capacity at a constant surface temperature of 65 °C with NH4OH at x1 = 0.3 and x1 = 0.5 – effect of liquid flow rate (a) and atomizing vapor pressure (b).
is the latent heat of vaporization, and V_ is the liquid flow rate supplied to the nozzle. Fig. 8 depicts the vaporization rate for selected NH4OH tests at x1 = 0.3 and x1 = 0.5 with the highest corresponding atomizing vapor pressures (179 kPa and 138 kPa) as a function of heat flux and liquid flow rate. Data indicate that the vaporization rate would be as high as 0.41 when the liquid flow rate is low at 1.05 ml/cm2.s and the heat flux is relatively high at 550 W/cm2. Higher liquid flow rates enable removing higher heat fluxes at comparable surface temperatures, although the coolant usage becomes less efficient as indicated by the lower vaporization ratio. Fig. 9 compares the data from NH4OH tests at x1 = 0.3 and x1 = 0.5 based on common liquid and vapor flow rate conditions. A major difference can be observed in the slope of the cooling curves that actually indicates the HTCs. Although the NH4OH mixture at x1 = 0.5 has a lower boiling point, at moderate-to-high heat flux levels (>300 W/cm2) NH4OH mixture at x1 = 0.3 with
Fig. 8. Vaporization ratio with NH4OH at x1 = 0.3 and x1 = 0.5 – effect of heat flux and liquid flow rate.
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Fig. 9. Comparison of spray cooling performance with NH4OH at x1 = 0.3 and x1 = 0.5 at common liquid flow rate and atomizing vapor pressure conditions.
2.10 ml/cm2.s liquid flow rate starts to provide a lower surface temperature. Fig. 10 shows the HTC (or h) values as a function of heat flux from the NH4OH tests at x1 = 0.3 and x1 = 0.5 included in Fig. 9. After excluding some of the HTC data at low heat fluxes, these data sets demonstrate clear trends, where NH4OH mixtures at x1 = 0.3 consistently provide higher HTC values and heat flux dependence compared to NH4OH mixtures at x1 = 0.5 for equivalent liquid and vapor flow rates. The level of HTCs, up to 125,000 W/m2°C in Fig. 10 (or 145,000 W/m2°C in this study corresponding to NH4OH test at x1 = 0.3 with 4.20 ml/cm2.s liquid flow rate and 179 kPa atomizing vapor pressure), is quite a bit lower compared to HTCs in some of the previous relevant work. Using air- or vapor-atomized nozzles, and plain, smooth heater surfaces, Yang et al. [2] and Chen et al. [3] reported HTC values of 250,000–500,000 W/m2 °C (at 500 W/cm2) with water (i.e., x1 = 0.0), and Bostanci et al. [5,6] obtained HTC values of 200,000–300,000 W/m2°C (at 500 W/cm2) with ammonia (i.e., x1 = 1.0). Although the heat transfer characteristics of two-phase spray cooling with binary mixtures have not been extensively studied and understood yet, a brief review of the current understanding would be helpful. Considering the spray conditions that form a 0.1–0.3 mm thick liquid film on the impinged surface [27] (including a much thinner, stagnant sub-layer region), major heat transfer mechanisms in
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spray cooling process involve; (i) convection, driven by the radial flow of liquid film, (ii) surface evaporation at the liquid-vapor interface, where heat transfer rate is governed by the temperature gradient across the liquid film, the thickness of the liquid film, and the thermophysical properties of the liquid, (iii) boiling through surface nucleation, maintained by active nucleation sites depending on the surface morphology, (iv) boiling through secondary nucleation, controlled by bubbles that are entrained as the liquid droplets enter the liquid film. Boiling therefore, plays a key role in the two-phase spray cooling process. In the single component boiling, the bubble growth is completely heat transfer controlled. However, in binary mixture boiling, mass transfer of the components may also affect the bubble growth. As outlined in [7], when the vaporization occurs in a non-azeotropic binary mixture (such as NH4OH), the vapor generated is richer in the more volatile component than the bulk liquid, and the remaining liquid in the vicinity of the interface has a correspondingly lower concentration of the more volatile component. Consequently, in the liquid phase, the more volatile component diffuses towards the interface, and the excess less volatile component diffuses away from the interface into the bulk liquid. As a result of this behavior, the boiling point at the liquid-vapor interface increases reflecting the local concentration level, and causes a reduction in the driving temperature difference for vaporization leading to depression in the boiling heat transfer rate. Such HTC depression is generally referred to as diffusion resistance to heat transfer. However, mass diffusion resistance in spray cooling could be significantly reduced since impinging droplets at bulk concentration constantly come in contact with liquid-vapor interface, and cause turbulent mixing and liquid replenishment. Based on the mixture theory [7], ideal HTC values can be determined by the reciprocal mass (or mole) fraction-weighted-average of the pure component HTCs at the specified heat flux and the same system pressure or temperature conditions. In an idealized two-phase heat transfer process with a non-azeotropic mixture (e.g., NH4OH), the HTC would vary from the highest value for the pure, less volatile component (e.g., water, x1 = 0.0), decrease as x1 increases (with HTC depression proportional to the difference between the liquid and vapor concentration levels), and reach the lowest value for the pure, more volatile component (e.g., ammonia, x1 = 1.0). Therefore, this HTC variation with x1 can explain the current NH4OH data where HTCs at x1 = 0.3 are higher than those at x1 = 0.5.
Fig. 10. Comparison of HTC with NH4OH at x1 = 0.3 and x1 = 0.5 at common liquid flow rate and atomizing vapor pressure conditions.
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4. Conclusions This study experimentally investigated the spray cooling performance of NH4OH binary mixtures at two selected ammonia mass fractions of, x1 = 0.3 and x1 = 0.5, representing near room temperature and near atmospheric pressure system conditions. Based on the obtained results the following conclusions can be reached: At a given atomizing vapor pressure (or vapor flow rate), increasing liquid flow rate between 1.05 and 4.20 ml/cm2.s increases the heat transfer rate (in a diminishing way, indicating an optimum level), due to higher droplet velocity and higher droplet flux. At a given liquid flow rate, increasing atomizing vapor pressure between 69 and 179 kPa also increases heat transfer rate, due to increased droplet velocity and decreased droplet size. HTC values of up to 145,000 W/m2°C from NH4OH mixtures are lower than those from water and ammonia. This HTC depression can be attributed to the mass diffusion resistance to heat transfer, an important phenomenon in boiling of binary mixtures, arising due to the difference between the liquid and vapor concentration levels during vaporization. However, the impact of diffusion resistance in spray cooling could be less important since impinging droplets at bulk concentration constantly come in contact with liquid-vapor interface, and cause turbulent mixing and liquid replenishment. HTCs decrease as x1 increases, following the expected trend for ideal HTC values based on mass (or mole) fraction-weightedaverage of the pure component HTCs. More comprehensive experimental studies, including full range of mass fractions (x1 = 0.0–1.0) are needed to gain additional insights and better understand the fundamentals of the twophase spray cooling with binary mixtures.
Acknowledgment We thank RINI Technologies Inc. for the financial support. References [1] L.C. Chow, M.S. Sehmbey, M.R. Pais, High heat flux spray cooling, in: C.-L. Tien (Ed.), Annual Review of Heat Transfer, vol. 8, Hemisphere Pub. Corp., New York, NY, 1997, pp. 291–318. [2] J. Yang, L.C. Chow, M.R. Pais, Nucleate boiling heat transfer in spray cooling, J. Heat Transfer 118 (1996) 668–671. [3] R.-H. Chen, L.C. Chow, J.E. Navedo, Effects of spray characteristics on critical heat flux in subcooled water spray cooling, Int. J. Heat Mass Transf. 45 (2002) 4033–4043. [4] R.-H. Chen, L.C. Chow, J.E. Navedo, Optimal spray characteristics in water spray cooling, Int. J. Heat Mass Transfer 47 (2004) 5095–5099.
[5] H. Bostanci, D.P. Rini, J.P. Kizito, V. Singh, S. Seal, L.C. Chow, High heat flux spray cooling with ammonia: investigation of enhanced surfaces for HTC, Int. J. Heat Mass Transf. 75 (2014) 718–725. [6] H. Bostanci, D.P. Rini, J.P. Kizito, V. Singh, S. Seal, L.C. Chow, High heat flux spray cooling with ammonia: investigation of enhanced surfaces for CHF, Int. J. Heat Mass Transf. 55 (2012) 3849–3856. [7] V.P. Carey, Liquid Vapor Phase Change Phenomena, 2nd ed., CRC Press, Boca Raton, FL, 2008. [8] J.R. Thome, S. Shakir, A new correlation for nucleate pol boiling of binary mixtures, AIChE Symp. Ser. 83 (1987) 46–51. [9] Y. Fujita, Q. Bai, M. Tsutsui, Heat transfer of binary mixtures in nucleate boiling, in: G.P. Celeta, P. DiMarco, A. Mariani (Eds.), 2nd Eur. Thermal Sci. and 14th UIT Nat. Heat Trans. Conf., 1996, pp. 1639–1646. [10] S.G. Kandlikar, Boiling heat transfer with binary mixtures: Part I – A theoretical model for pool boiling, J. Heat Transfer 120 (1998) 380–387. [11] W.R. McGillis, V.P. Carey, On the role of Marangoni effects on the critical heat flux for pool boiling of binary mixtures, J. Heat Transfer 118 (1996) 103–109. [12] Y. Fujita, Q. Bai, Critical heat flux of binary mixtures in pool boiling and its correlation in terms of Marangoni number, Int. J. Refrig 20 (1998) 616–622. [13] V.V. Yagov, Critical heat prediction for pool boiling of binary mixtures, Chem. Eng. Res. Des. 82 (2004) 457–461. [14] B. Abbasi, J. Kim, Prediction of PF-5060 spray cooling heat transfer and critical heat flux, J. Heat Transfer 133 (2011) 101504–101515. [15] E.A. Silk, J. Kim, K. Kiger, Spray cooling of enhanced surfaces: Impact of structured surface geometry and spray axis inclination, Int. J. Heat Mass Transf. 49 (2006) 4910–4920. [16] M. Visaria, I. Mudawar, Application of two-phase spray cooling for thermal management of electronic devices, IEEE Trans. Compon. Packag. Technol. 32 (2009) 784–793. [17] A.G. Pautsch, T.A. Shedd, Spray impingement cooling with single and multiplenozzle arrays. Part I: Heat transfer data using FC-72, Int. J. Heat Mass Transf. 48 (2005) 3167–3175. [18] L. Lin, R. Harris, J. Lawson, R. Ponnappan, Spray cooling with methanol and water mixtures, in: Proceedings of 9th AIAA/ASME Joint Thermophysics and Heat Transfer Conference, San Francisco, CA, 2006. [19] L.J. Turek, D.P. Rini, B.A. Saarloos, L.C. Chow, Evaporative spray cooling of power electronics using high temperature coolant, in: Proceedings of ITHERM 2008, Orlando, FL, 2008. [20] L.J. Turek, D.P. Rini, B.A. Saarloos, L.C. Chow, Enabling much higher power densities in aerospace power electronics with high temperature evaporative spray cooling, in: Proceedings of SAE Power Systems Conference, Seattle, WA, 2008. [21] H. Bostanci, D. Van Ee, B.A. Saarloos, D.P. Rini, L.C. Chow, Thermal management of power inverter modules at high heat fluxes via two-phase spray cooling, IEEE Trans. Compon. Packag. Manuf. Technol. 2 (2012) 1480–1485. [22] P.N. Karpov, A.D. Nazarov, A.F. Serov, V.I. Terekhov, Evaporative cooling by a pulsed jet spray of binary ethanol-water mixture, Tech. Phys. Lett. 41 (2015) 668–671. [23] M.R. Conde-Petite, Thermophysical Properties of NH3 + H2O Solutions for the Industrial Design of Absorption Refrigeration Equipment, Technical Report, M. Conde Engineering, Zurich, Switzerland, 2004, pp. 1–34. [24] B.H. Jennings, F.P. Shannon, The thermodynamics of absorption refrigeration – Part I Tables of the properties of aqua-ammonia solutions, Refrig. Eng. (5) (1938). [25] DuPont Technical Information P134, DuPont HFC-134a properties, uses, storage, and handling, 2004.. [26] 3M Product Information, 3M FluorinertTM Electronic Liquid FC-72, 2000. . [27] J. Yang, M.R. Pais, L.C. Chow, A. Ito, Liquid film thickness and topography determination using Fresnel diffraction and holography, Exp. Heat Transfer 5 (1992) 239–252.
Contents lists available at ScienceDirect
International Journal of Heat and Mass Transfer journal homepage: www.elsevier.com/locate/ijhmt
Spray cooling with ammonium hydroxide Huseyin Bostanci ⇑, Bin He, Louis C. Chow Department of Mechanical and Aerospace Engineering, University of Central Florida, Orlando, FL 32816, United States
a r t i c l e
i n f o
Article history: Received 26 April 2016 Received in revised form 8 November 2016 Accepted 10 November 2016
Keywords: Thermal management Spray cooling High heat flux Binary mixture Ammonium hydroxide Diffusion resistance
a b s t r a c t An experimental study was conducted to investigate the performance characteristics of spray cooling with ammonium hydroxide (NH4OH) binary mixture for high heat flux removal. Two mixtures, having ammonia mass fractions of 0.3 and 0.5, were selected to represent practical operation conditions near atmospheric pressure and room temperature. Experimental setup involved a closed loop system with a vapor atomized spray nozzle and a 1-cm2 heater sample that simulated a high heat flux source. Tests were performed with gradually increasing heat fluxes of up to 800 W/cm2 and maintaining surface temperatures below 75 °C at varying liquid and vapor flow rates. Results indicated that the heat transfer coefficient (HTC) values from NH4OH mixtures can be lower than those from pure water and pure ammonia. The data suggested that boiling depression, due to mass diffusion resistance at liquid vapor interface, could greatly affect the overall spray cooling performance, especially when the binary mixtures comprise components with widely different boiling points. The study therefore provides performance characteristics, as well as some fundamental insights, for a potential spray cooling scheme suitable for low temperature, low pressure operations in various applications including thermal management of aerospace electronics and electro-optics. Ó 2016 Elsevier Ltd. All rights reserved.
1. Introduction Two-phase spray cooling has been an emerging thermal management technique characterized by its major capabilities of high heat transfer rate, near-uniform surface temperature, and efficient coolant usage that leads to compact and lightweight systems. Due to these capabilities, spray cooling is a promising approach for high heat flux applications in computing, power electronics, and electro-optics. Selection of working fluid is very important as it directly affects both performance characteristics and operation conditions of a cooling system. The working fluid should have a high latent heat of vaporization to remove high heat flux levels, a boiling point close to the desired operating temperature of the target device being cooled, and a freezing point low enough to provide protection against system freeze-up at anticipated cold ambient. A proper temperature-pressure relationship is also critical to avoid excessively low pressure (high vacuum) or excessively high pressure in the system. Moreover, the working fluid should be chemically stable, noncorrosive, and safe.
⇑ Corresponding author at: University of North Texas, Department of Engineering Technology, UNT Discovery Park, 3940 North Elm St. F115, Denton, TX 76207, United States. E-mail address: [email protected] (H. Bostanci). http://dx.doi.org/10.1016/j.ijheatmasstransfer.2016.11.035 0017-9310/Ó 2016 Elsevier Ltd. All rights reserved.
Certain applications, such as thermal management of highpower diode-pumped solid-state lasers for aerospace systems, poses multiple challenges and necessitates careful consideration to determine a proper working fluid. The laser diode arrays generate high heat fluxes at substrate-level (>600 W/cm2), they mandate operation at low temperatures (<25 °C) to attain high electrical-tooptical energy conversion efficiency, and the emitters (and surface temperature of diode arrays) also require nearly-uniform temperature distributions, within a few °C, to achieve high quality optical beam. Due to the ambient conditions for aerospace applications, the cooling system should be able to operate at very low temperatures (70 °C) without freezing issues. Additionally, the system pressure should ideally be close to atmospheric pressure levels during operation. Based on all these challenging conditions, two-phase spray cooling seems to be the appropriate thermal management scheme. Considering thermophysical properties, particularly high latent heat of vaporization, water and ammonia are the two best heat transfer fluids. However, each fluid has undesirable saturation temperature-pressure characteristics for diode laser cooling. In order to use water for low temperature cooling, its boiling point should be reduced significantly from 100 °C at 1 bar to 20 °C at 0.02 bar by maintaining the system under high vacuum. This would then complicate the system design, and cause potential leak issues. Water also has a very high freezing point for aerospace con-
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Nomenclature A CHF h hfg HTC I k l
area, cm2 critical heat flux, W/cm2 heat transfer coefficient, kW/m2 °C latent heat of vaporization, kJ/kg heat transfer coefficient, kW/m2 °C current, A heater wall thermal conductivity, W/m °C TC to spray surface distance in heater wall, m
ditions. Ammonia appears to be an ideal working fluid for spray cooling of diode laser arrays due to its high latent heat, low boiling point and low freezing point at the normal pressure. However, ammonia has a very strong temperature-pressure dependency. The ammonia pressure can easily exceed 10 bar at room temperature. High saturation pressure (4–10 bars) would deform the heat source structure (i.e., diode bar package) and may lead to failure. This is especially true when the substrate accommodating the heat source is made very thin in order to minimize the internal thermal resistance. Therefore, for room temperature spray cooling applications, water dictates very low system pressure leading to serious fluid sealing problems, while ammonia dictates very high system pressure leading to serious structural problems. In this case, utilizing a mixture of compatible fluids is an attractive approach to obtain the desirable saturation temperature-pressure behavior needed for spray cooling of diode lasers. Ammonium hydroxide (NH4OH) is made by dissolving ammonia into water. Varying the mass fraction of the solution can modify the saturation temperature-pressure relationship. For instance, NH4OH with an ammonia mass fraction, x1, of 0.45 provides 0 °C saturation temperature at approximately 1 bar pressure. Its latent heat of vaporization is still 5–10 higher than those of common refrigerants. The low freezing point (around 80 °C for x1 >0.3) also prevents the plumbing from freezing when spray systems are used in aerospace vehicles. This binary fluid takes advantage of the high latent heat of vaporization associated with the parent fluids without the drawbacks of excessively low or high saturation pressures of water and ammonia, respectively. In addition, NH4OH has a reasonably high thermal conductivity comparable to that of water. Both the parent fluids of NH4OH have very favorable spray cooling heat transfer rates as reported in the earlier research [1–6]. If the binary mixture of NH4OH can provide cooling characteristics somewhere between those of its parent fluids – water and ammonia, it would enable spray cooling technology for high heat flux cooling capability at room temperature operation. This would then allow effective cooling of the delicate electronics and optics devices, such as diode laser arrays, without the excessively high or low system pressure. Binary mixtures have been studied extensively for boiling heat transfer due to their wide use in industrial applications, such as petro-chemical, refrigeration, and power generation. Carey [7] provided a comprehensive review on the pool boiling of binary mixtures addressing thermodynamics aspects and major heat transfer characteristics, specifically the effects of mass diffusion resistance on nucleate boiling, and surface tension gradients on CHF. Some of the exemplary efforts in this area focused on developing prediction methods, including Thome and Shakir [8], Fujita et al. [9], and Kandlikar [10] regarding HTC, and McGillis and Carey [11], Fujita and Bai [12], and Yagov [13] regarding critical heat flux (CHF) correlations. However, there are very limited spray cooling research focusing on binary mixtures, as previous efforts mainly
q00 Tsurf TCavg V V_ V_ req x1
q
heat flux, W/cm2 surface temperature, °C average thermocouple reading, °C voltage, V supplied liquid flow rate (ml/cm2.s) required liquid flow rate (ml/cm2.s) liquid mass fraction of more volatile (NH3) component liquid density (kg/m3)
considered single-component working fluids [1–6,14–17], and none of which, to the best of authors’ knowledge, studied NH4OH. Lin et al. [18] investigated spray cooling with water/methanol mixtures at methanol concentrations of 0, 20, 50 and 100% by volume. Their experimental conditions involved various chamber pressures to maintain saturation temperatures of 45 and 65 °C, and relatively low heat fluxes in the range of 15–75 W/cm2. The data showed that HTCs were highest for water, and lowest for methanol. The performance of water/methanol mixtures was found to be between those of pure water and methanol, although in some conditions, the mixture with 50% ethanol concentration offered comparable performance to that of pure water. The authors noted that the nucleate boiling is a major heat transfer mode in spray cooling, and flashing of the more volatile component at the nozzle exit helps with atomizing liquid droplets and enhance the heat transfer performance. Turek et al. [19] evaluated the capabilities of spray cooling for power inverter applications using water/ propylene glycol (WPG) at a 50% concentration by volume. They replaced the single-phase liquid cooled base plate of a commercial power inverter module with a custom-made base plate of comparable size, and integrated a spray nozzle array. DBC board in the module was sprayed with pressure atomizer nozzles featuring 0.19 l/cm2min liquid flow rate and 275 kPa pressure drop. Using WPG at 100 °C, heat fluxes of up to 148 W/cm2 were removed without exceeding Insulated Gate Bipolar Transistor (IGBT) junction temperature of 125 °C. Their spray cooling approach reduced the overall thermal resistance with the help of a very high HTC and achieved up to a 3.4 increase in inverter power level. Later, Turek et al. [20] conducted experiments with pure water at 95 °C using the same setup. IGBT junction temperatures for the water spray cooling was about 10 °C lower compared to WPG spray cooling, although the difference in the boiling points of two cases was 6 °C. Considering the similar subcooling levels, they concluded that the vapor generated by WPG boiling is nearly all water, leaving a higher concentration mixture at the spray surface with a higher boiling. This locally higher boiling point diminishes the boiling performance and leads to higher temperatures. Bostanci et al. [21] evaluated a two-phase spray cooling scheme for the thermal management of automotive power inverter modules using an antifreeze coolant (water/alcohol mixture). The experimental setup included pressure atomized spray nozzles with 0.15 l/cm2min liquid flow rate and 145 kPa pressure drop, and featured two types of enhanced spray surface with microscale structures. The spray cooling tests with azeotropic mixture at 88 °C saturation temperature and atmospheric pressure, showed that up to 400 W/cm2 heat fluxes can be reached with only 14 °C surface superheat, resulting in a HTC of 280,000 W/m2 °C. Compared to single-phase convective cooling, this spray cooling system offered up to 5.5 potential increase in inverter power. Most recently, Karpov et al. [22] performed an experimental study with water/ethanol mixtures, at ethanol concentrations of 0–96% by mass, and used air atomized
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47
spray nozzle arrays in a pulsed mode resulting in very low flow rates of up to 0.0173 g/cm2.s. They examined the effects of spray duty cycle and frequency with an isothermal heater at 70 °C, and observed increasing HTC values with increasing ethanol concentration (with a max. at 50–60%). The higher HTC was explained with the accelerated evaporation rates driven by the ethanol content. This study aims to experimentally investigate the performance characteristics of spray cooling with NH4OH binary mixture for high heat flux removal at practical operation conditions near atmospheric pressure and room temperature. The initial data will help understand the heat transfer mechanisms for binary mixture spray cooling, and be instrumental in the development of a high heat flux cooling technique for low temperature, low pressure operations in various applications including thermal management of aerospace electronics and electro-optics. 2. Experimental setup and procedure 2.1. Working fluid Ammonium hydroxide (NH4OH) was selected as the working fluid in this study. Aqua ammonia, aqueous ammonia and ammonium hydroxide are synonymous terms referring to a solution of ammonia in water. Achieving a high heat flux cooling technique at low temperature and pressure operation conditions requires to use a working fluid with proper saturation temperature-pressure relations. Fig. 1 includes such saturation temperature-pressure values for water, ammonia and two other NH4OH mixtures. Defining the desired system conditions to be 1.0–1.25 bar pressure, and 0– 30 °C temperature (shown with shaded areas in the plot), NH4OH mixtures at ammonia mass fractions (x1) of 0.3 and 0.5 emerge as proper working fluids. Specifically, NH4OH at x1 = 0.3 has a boiling point of 27 °C under 1 bar, (that would allow implementing thin structures with low thermal resistance due to pressure level), and NH4OH at x1 = 0.5 has a boiling point of 0 °C under about 1.25 bar (that would allow high electrical-to-optical energy conversion efficiencies due to low temperature level). These two NH4OH mixtures would also provide good freezing protection. Fig. 2a shows the freezing point (along with boiling point) of NH4OH as a function of x1 at atmospheric pressure, where NH4OH at
Fig. 2. Boiling and freezing points (a), and latent heat of vaporization (b) of NH4OH as a function of x1 (based on [23,24]).
x1 = 0.3 and x1 = 0.5 have a freezing point of 680 °C. Fig. 2b includes latent heat of vaporization for the NH4OH mixtures, where it sharply drops beyond x1 = 0.1. Table 1 summarizes the critical thermophysical properties of the selected NH4OH binary mixtures, their parent (pure component) fluids, and some other common refrigerants. When compared to HFC-134a and FC-72, NH4OH at x1 = 0.3 and x1 = 0.5 possess much higher latent heat (7–18) and thermal conductivity (6–10) values. Based on the outlined considerations, NH4OH at x1 = 0.3 and x1 = 0.5 were selected for the present study. While the NH4OH at x1 = 0.3 is commercially available, NH4OH at x1 = 0.5 is not commonly offered by chemical vendors. Therefore, NH4OH at x1 = 0.5 was prepared on site by pulling vacuum on an empty storage cylinder, and then carefully filling it with equal masses of distilled water and anhydrous ammonia. NH4OH is corrosive to copper, copper containing alloys (such as brass), zinc, cadmium, silver, aluminum alloys and galvanized surfaces. Thus, the use of NH4OH as the working fluid requires special attention in materials selection during the system design. Furthermore, such cooling systems should be sealed well and located in non-confined areas, as NH4OH leaks would pose health hazards. 2.2. Spray cooling system
Fig. 1. Saturation temperature and pressure relationship for water, ammonia, and NH4OH mixtures (based on [23]).
A closed loop spray cooling system was used to run the experiments. Schematic diagram in Fig. 3 illustrates the three main sections of the system: NH4OH liquid loop, NH4OH vapor loop, and HFC-134a loop. A reservoir, measuring 10 cm in diameter and 19 cm in length, forms the center of the system containing liquid and vapor phases
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Table 1 Thermophysical properties of selected working fluids (based on [23–26]). Properties at 1.013 bar
Water
Ammonia
NH4OH (x1 = 0.3)
NH4OH (x1 = 0.5)
FC-72
HFC-134a
Freezing point (°C) Boiling point (°C) Latent heat of vaporization (kJ/kg) Thermal conductivity (W/m °C) at 25 °C
0 100 2257 0.614
77.7 33.3 1371 0.466
82.7 27.2 1474 0.548
80.1 4.2 1560 0.526
90 56 88 0.057
103.3 26.1 217.2 0.082
and vapor loops were placed inside a fume hood to ensure the NH3 concentration was always at safe levels (<200 ppm) even in the case of leaks, or during the maintenance of the setup. All wetted parts of the system featured NH4OH compatible materials including stainless steel (chamber), PTFE (tubing), polypropylene (fittings), and acrylic (chamber window). The test setup was equipped with thermocouple (TC) probes, pressure transducers, and a computer controlled data acquisition system for accurate data recording. 2.3. Spray nozzles
Fig. 3. Schematic diagram of the spray cooling setup.
of NH4OH, a test heater, and spray nozzles (a main nozzle for cooling the test heater, and four auxiliary nozzles for condensing the generated vapor). NH4OH liquid loop involves a gear pump that takes saturated NH4OH liquid from the bottom of the reservoir and provides a pressure increase. Part of this liquid flow was supplied to the spray nozzle at a controlled rate. The rest of the liquid flow was routed to a heat exchanger to lower the liquid temperature, and the resulting subcooled liquid was then distributed to the four auxiliary nozzles mounted symmetrically onto the reservoir wall near the top to condense the generated vapor and control the spray chamber pressure/temperature. Use of these side sprays as the condenser allowed a greatly expanded heat exchange area involving surface area of the spray droplets, and helped maintain the mixture concentration level in the liquid pool. NH4OH vapor loop involves a diaphragm compressor that takes saturated NH4OH vapor from the upper section of the reservoir and provides the required pressure lift. The vapor flow at the adjusted rate is then supplied to the main nozzle, where it mixes with the liquid stream and facilitates spray atomization. Additionally, the experimental setup has a HFC-134a (a hydrochlorofluorocarbon compound) refrigerant loop that interacts with the NH4OH liquid loop through the heat exchanger. This loop is a typical vapor compression cycle that absorbs heat from NH4OH liquid loop and rejects it to the circulating tap water. Spray chamber temperature and pressure was controlled in two ways: (i) by adjusting the NH4OH liquid flow rate that goes into the auxiliary nozzles, and (ii) by adjusting the temperature of the heat exchanger through low side pressure of the HFC-134a loop. The reservoir also had a port for charging/ discharging the working fluid. The reservoir, and the NH4OH liquid
A compact vapor atomized spray nozzle (Fig. 4a), manufactured by RINI Technologies, Inc., was used as the main nozzle for cooling the test heater. In these nozzles a fine liquid stream is injected into a high velocity vapor stream. The shear force created by the vapor stream atomizes the liquid into fine droplets that are ejected from the nozzle orifice. The nozzle produces a full-cone spray with an angle of 50° that requires a nozzle-to-surface distance of 10 mm to adequately cover the spray surface. The uniformity of droplet density for this nozzle type was found to be within 20% in radial direction from center to edge. The pressure and temperature of the liquid and vapor supplies, as well as the spray chamber were measured to determine the driving pressures across the nozzle, and saturation condition of the coolant. The other four auxiliary nozzles, used for condensing the vapor, were pressure atomized nozzles featuring a simpler design that requires only liquid supply. All the nozzles were made out of stainless steel. 2.4. Heater and spray surface The test heater design is shown in Fig. 4b–c. One side of the heater had a 1 cm 1 cm protrusion (Fig. 4b) where a matching size (1-cm2) thick film resistor (a flip chip type resistor with backpad metallization) was soldered to simulate a heat generating device. In an actual cooling system configuration, a heat-generating device is usually attached to a substrate or an interface layer that is exposed to a spray/cooling medium, and the interfacial thermal contact resistance is inevitable. Thus, it is very important to minimize the resistance for low impact on the device temperature. In this study, the resistor was carefully soldered onto the heater to ensure that the solder layer was thin and free of air pockets. In terms of temperature measurements in the experiments, the contact resistance has no impact as the measurements are away from the interface. The opposite side of the heater featured a similar size (0.9 cm 0.9 cm) recessed area (Fig. 4c) resulting in a 2 mm thick copper substrate under the heat generating resistor, and it was exposed to spray from the main nozzle. The heater was made of copper (C101 grade), a material not compatible with ammonia. In order to provide corrosion protection, spray side of the test heater was coated with a thin (40–60 lm) layer of tin-lead solder resulting in a relatively smooth surface. The heat flux, q00 , was determined from the total power supplied into the thick film resistor per unit area, and it is defined as q00 = V I/A, where V is the DC voltage applied to the resistor, I is the resulting current for 10 Ohm resistance, and A is the resistor base. The heater temperature was monitored with two type-T TCs
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Fig. 4. Vapor atomized spray nozzle (a), and test section (heater) resistor side (b) and spray side (c).
embedded at the mid-section in the heater wall (1 mm below the surface). One of the TCs was located at the center, while the other TC was located 1 mm away from the center as illustrated in Fig. 4b. Temperature difference between the two thermocouple readings did not exceed ±1 °C at 200 W/cm2, and ±2.0 °C at 500 W/cm2 indicating uniform cooling across the surface. The spray surface temperature, Tsurf, was calculated with Fourier’s 1-D steady-state heat conduction equation, and is defined as Tsurf = TCavg (q00 l)/k, where TCavg is the average of two TC readings, l is the distance from TC location to the surface (taken as 1 mm for copper substrate and 50 lm for tin-lead coating layer), and k is the thermal conductivity of the copper and tin-lead solder. 2.5. Test conditions and procedure Spray cooling setup was maintained at saturation conditions of approximately 27 °C temperature and 1 bar pressure for NH4OH tests at x1 = 0.3, and approximately 0 °C and 1.25 bar pressure for NH4OH tests at x1 = 0.5. Cooling curves, in the form of surface temperature (Tsurf) vs. heat flux (q00 ), were generated for each test by increasing the heat flux gradually in steps of 45–65 W/cm2 until the surface temperature reaches approximately 75 °C. This temperature limit was determined considering practical operation of solid state lasers, which rely on low temperatures for achieving high conversion efficiency. At each heat flux level, the resulting heater temperatures were recorded every 2 s over 3 min durations. Major experimental parameters in this study were the liquid flow rate and atomizing vapor pressure (or vapor pressure drop across the nozzle that is proportional to vapor flow rate) to investigate their effects on the spray cooling heat transfer performance. For NH4OH tests at x1 = 0.3, liquid flow rates (per unit area) and atomizing vapor pressures were varied between 1.05–4.20 ml/cm2.s, and 69–179 kPa (10–26 psi), respectively. For NH4OH tests at x1 = 0.5, liquid flow rates and atomizing vapor pressures were selected to be 1.05–2.10 ml/cm2.s, and 69–138 kPa (10–20 psi), respectively. These parameters resulted in a total of 14 experiments as described in Table 2.
Table 2 Test descriptions. Test ID
NH4OH, x1
Liquid flow rate (ml/cm2.s)
Atomizing vapor pressure (kPa)
1 2 3 4 5 6 7 8 9 10
0.3
1.05 2.10 3.15 1.05 2.10 3.15 1.05 2.10 3.15 4.20
69
11 12 13 14
0.5
1.05 2.10 1.05 2.10
138
179
69 138
film resistor to heater body via conduction was estimated to be approximately 4% based on finite element analysis results. Heat loss from the thick film resistor to the ambient environment was negligibly small (<1 W) based on calculations considering natural convection and black body radiation from 100 °C heater surface to 20 °C stagnant air.
3. Results and discussion Tests described in Table 2 were conducted to investigate the spray cooling performance with NH4OH binary mixture at two concentration levels and at varying liquid and vapor flow rates. Fig. 5
2.6. Uncertainty analysis Experimental uncertainties were estimated for all measurements that are critical in performance evaluation. Error involved in the heat flux measurement (considering variations in voltage, current, and area) was ±1.3% at 500 W/cm2. Error in the temperature measurements from the embedded TCs in the heater wall was calibrated to be ±0.2 °C. Spray surface temperature had an uncertainty of approximately ±0.5 °C involving uncertainty in temperature extrapolation from the TC hole to the heater surface at 500 W/cm2. The HTC calculation included ±3.4% uncertainty at 500 W/cm2. Flow rate measurements had ±5% uncertainty due to variable area flow meter characteristics. Heat loss from the thick
Fig. 5. Spray cooling performance with NH4OH at x1 = 0.3 – effect of liquid flow rate and atomizing vapor pressure.
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includes all data from NH4OH tests at x1 = 0.3 representing 27 °C saturation temperature at 1 bar pressure. The change in saturation temperature during these tests (at heat fluxes of >0 W/cm2), due to fluctuations in chamber pressure, was maintained within 3–4 °C range. Heat fluxes were increased gradually until surface temperature reaches up to 75 °C, and the highest heat flux (but, not CHF) was 800 W/cm2 at 3.15–4.20 ml/cm2.s liquid flow rate and 179 kPa atomizing vapor pressure. Each data set was fitted with a second order polynomial curve to better represent and compare the heat transfer performance. When the effect of liquid flow rate (at a given atomizing vapor pressure) is examined (corresponding to same colored data sets with varying shapes), it is obvious that increasing liquid flow rate provides a better heat transfer, as expected, since higher liquid flow rate would lead to higher droplet velocity and higher droplet flux. However, the level of improvement diminishes beyond 2.10 ml/cm2.s suggesting an optimum level. This observation is consistent for atomizing vapor pressures at 69, 138, and 179 kPa. When the effect of atomizing vapor pressure (at a given liquid flow rate) is considered (corresponding to data sets with similar shapes and different colors), increasing atomizing vapor pressure also improves cooling performance, again as expected, since higher atomizing vapor pressure would lead to higher droplet velocity, and smaller droplet sizes. This trend is consistent at 1.05, 2.10, and 3.15 ml/cm2.s liquid flow rates. Fig. 6 presents data from the NH4OH tests at x1 = 0.5 representing 0 °C saturation temperature at 1.25 bar pressure. Similarly, the change in saturation temperature, due to fluctuations in chamber pressure, was limited to 3–4 °C range. The highest heat flux attained at a surface temperature of 75 °C was approximately 500 W/cm2. Increasing the liquid flow rate at a given atomizing vapor pressure at this concentration level resulted in nearly the same cooling performance. Increasing the atomizing vapor pressure at a given liquid flow rate however, helped improve the cooling performance. The effects of liquid flow rate and atomizing vapor pressure on cooling performance are also illustrated in Fig. 7 to offer an easier comparison and specify the heat removal capability at a fixed 65 °C surface temperature for various conditions with both NH4OH mixtures at x1 = 0.3 and x1 = 0.5. Since the liquid and vapor streams are supplied in saturated condition, the heat absorption on the spray surface is attained mainly by the latent heat during liquid vapor phase change. In this case, vaporization rate can be estimated by calculating the required liquid flow rate to absorb a given heat flux over the actual liquid flow rate used, and it can be defined as _ where q is the NH4OH liquid density, hfg V_ req =V_ ¼ ðq00 =ðq hfg ÞÞ=V,
Fig. 6. Spray cooling performance with NH4OH at x1 = 0.5 – effect of liquid flow rate and atomizing vapor pressure.
Fig. 7. Heat removal capacity at a constant surface temperature of 65 °C with NH4OH at x1 = 0.3 and x1 = 0.5 – effect of liquid flow rate (a) and atomizing vapor pressure (b).
is the latent heat of vaporization, and V_ is the liquid flow rate supplied to the nozzle. Fig. 8 depicts the vaporization rate for selected NH4OH tests at x1 = 0.3 and x1 = 0.5 with the highest corresponding atomizing vapor pressures (179 kPa and 138 kPa) as a function of heat flux and liquid flow rate. Data indicate that the vaporization rate would be as high as 0.41 when the liquid flow rate is low at 1.05 ml/cm2.s and the heat flux is relatively high at 550 W/cm2. Higher liquid flow rates enable removing higher heat fluxes at comparable surface temperatures, although the coolant usage becomes less efficient as indicated by the lower vaporization ratio. Fig. 9 compares the data from NH4OH tests at x1 = 0.3 and x1 = 0.5 based on common liquid and vapor flow rate conditions. A major difference can be observed in the slope of the cooling curves that actually indicates the HTCs. Although the NH4OH mixture at x1 = 0.5 has a lower boiling point, at moderate-to-high heat flux levels (>300 W/cm2) NH4OH mixture at x1 = 0.3 with
Fig. 8. Vaporization ratio with NH4OH at x1 = 0.3 and x1 = 0.5 – effect of heat flux and liquid flow rate.
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Fig. 9. Comparison of spray cooling performance with NH4OH at x1 = 0.3 and x1 = 0.5 at common liquid flow rate and atomizing vapor pressure conditions.
2.10 ml/cm2.s liquid flow rate starts to provide a lower surface temperature. Fig. 10 shows the HTC (or h) values as a function of heat flux from the NH4OH tests at x1 = 0.3 and x1 = 0.5 included in Fig. 9. After excluding some of the HTC data at low heat fluxes, these data sets demonstrate clear trends, where NH4OH mixtures at x1 = 0.3 consistently provide higher HTC values and heat flux dependence compared to NH4OH mixtures at x1 = 0.5 for equivalent liquid and vapor flow rates. The level of HTCs, up to 125,000 W/m2°C in Fig. 10 (or 145,000 W/m2°C in this study corresponding to NH4OH test at x1 = 0.3 with 4.20 ml/cm2.s liquid flow rate and 179 kPa atomizing vapor pressure), is quite a bit lower compared to HTCs in some of the previous relevant work. Using air- or vapor-atomized nozzles, and plain, smooth heater surfaces, Yang et al. [2] and Chen et al. [3] reported HTC values of 250,000–500,000 W/m2 °C (at 500 W/cm2) with water (i.e., x1 = 0.0), and Bostanci et al. [5,6] obtained HTC values of 200,000–300,000 W/m2°C (at 500 W/cm2) with ammonia (i.e., x1 = 1.0). Although the heat transfer characteristics of two-phase spray cooling with binary mixtures have not been extensively studied and understood yet, a brief review of the current understanding would be helpful. Considering the spray conditions that form a 0.1–0.3 mm thick liquid film on the impinged surface [27] (including a much thinner, stagnant sub-layer region), major heat transfer mechanisms in
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spray cooling process involve; (i) convection, driven by the radial flow of liquid film, (ii) surface evaporation at the liquid-vapor interface, where heat transfer rate is governed by the temperature gradient across the liquid film, the thickness of the liquid film, and the thermophysical properties of the liquid, (iii) boiling through surface nucleation, maintained by active nucleation sites depending on the surface morphology, (iv) boiling through secondary nucleation, controlled by bubbles that are entrained as the liquid droplets enter the liquid film. Boiling therefore, plays a key role in the two-phase spray cooling process. In the single component boiling, the bubble growth is completely heat transfer controlled. However, in binary mixture boiling, mass transfer of the components may also affect the bubble growth. As outlined in [7], when the vaporization occurs in a non-azeotropic binary mixture (such as NH4OH), the vapor generated is richer in the more volatile component than the bulk liquid, and the remaining liquid in the vicinity of the interface has a correspondingly lower concentration of the more volatile component. Consequently, in the liquid phase, the more volatile component diffuses towards the interface, and the excess less volatile component diffuses away from the interface into the bulk liquid. As a result of this behavior, the boiling point at the liquid-vapor interface increases reflecting the local concentration level, and causes a reduction in the driving temperature difference for vaporization leading to depression in the boiling heat transfer rate. Such HTC depression is generally referred to as diffusion resistance to heat transfer. However, mass diffusion resistance in spray cooling could be significantly reduced since impinging droplets at bulk concentration constantly come in contact with liquid-vapor interface, and cause turbulent mixing and liquid replenishment. Based on the mixture theory [7], ideal HTC values can be determined by the reciprocal mass (or mole) fraction-weighted-average of the pure component HTCs at the specified heat flux and the same system pressure or temperature conditions. In an idealized two-phase heat transfer process with a non-azeotropic mixture (e.g., NH4OH), the HTC would vary from the highest value for the pure, less volatile component (e.g., water, x1 = 0.0), decrease as x1 increases (with HTC depression proportional to the difference between the liquid and vapor concentration levels), and reach the lowest value for the pure, more volatile component (e.g., ammonia, x1 = 1.0). Therefore, this HTC variation with x1 can explain the current NH4OH data where HTCs at x1 = 0.3 are higher than those at x1 = 0.5.
Fig. 10. Comparison of HTC with NH4OH at x1 = 0.3 and x1 = 0.5 at common liquid flow rate and atomizing vapor pressure conditions.
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4. Conclusions This study experimentally investigated the spray cooling performance of NH4OH binary mixtures at two selected ammonia mass fractions of, x1 = 0.3 and x1 = 0.5, representing near room temperature and near atmospheric pressure system conditions. Based on the obtained results the following conclusions can be reached: At a given atomizing vapor pressure (or vapor flow rate), increasing liquid flow rate between 1.05 and 4.20 ml/cm2.s increases the heat transfer rate (in a diminishing way, indicating an optimum level), due to higher droplet velocity and higher droplet flux. At a given liquid flow rate, increasing atomizing vapor pressure between 69 and 179 kPa also increases heat transfer rate, due to increased droplet velocity and decreased droplet size. HTC values of up to 145,000 W/m2°C from NH4OH mixtures are lower than those from water and ammonia. This HTC depression can be attributed to the mass diffusion resistance to heat transfer, an important phenomenon in boiling of binary mixtures, arising due to the difference between the liquid and vapor concentration levels during vaporization. However, the impact of diffusion resistance in spray cooling could be less important since impinging droplets at bulk concentration constantly come in contact with liquid-vapor interface, and cause turbulent mixing and liquid replenishment. HTCs decrease as x1 increases, following the expected trend for ideal HTC values based on mass (or mole) fraction-weightedaverage of the pure component HTCs. More comprehensive experimental studies, including full range of mass fractions (x1 = 0.0–1.0) are needed to gain additional insights and better understand the fundamentals of the twophase spray cooling with binary mixtures.
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